Encyclopedia of Chemical Processing and Design. 69_Supplement 1

Encyclopedia of Chemical Processing and Design: 69 Supplement 1

Rayford G. Anthony John J. McKetta

Marcel Dekker, Inc.

Encyclopedia of Chemical Processing and Design EDITOR

Rayford G. Anthony

SENIOR ADVISORY EDITOR

John J. McKetta

69

Supplement 1

Library of Congress Cataloging in Publication Data Main entry under title: Encyclopedia of chemical processing and design. Includes bibliographic references. 1. Chemical engineering—Dictionaries 2. Technical—Dictionaries. I. McKetta, John J. II. Cunningham, William Aaron. Tp9.E66 660.2′8′003 ISBN: 0-8247-2621-9

Chemistry, 75-40646

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Contributors to Volume 69 Steve Chum, Ph.D. Research Fellow, The Dow Chemical Company, Polyolefins Research, Freeport, Texas: Structure, Properties, and Applications of Polyolefins Produced by Single-Site Catalyst Technology Ray A. Cocco, Ph.D. Senior Specialist, The Dow Chemical Company, Midland, Michigan: Circulating Fluidized Bed Reactors: Basic Concepts and Hydrodynamics David M. Fishbach, P.E. Senior Consulting Engineer, Starfire Electronic Development & Marketing, Ltd., Bloomfield Hills, Michigan: Nanophase Materials in Chemical Process Avery N. Goldstein, Ph.D. Research Director, Starfire Electronics Development & Marketing, Ltd., Bloomfield Hills, Michigan: Nanophase Materials in Chemical Process Manfred Grove Senior Partner, Intermacom A.G., Technology Consultants, Auerich, Switzerland: Introduction to the Selective Catalytic Reduction Technology Dennis Hendershot Rohm and Haas Company, Bristol, Pennsylvania: Fundamentals of Process Safety and Risk Management Trevor A. Kletz Process Safety Consultant, Cheshire, United Kingdom: Fundamentals of Process Safety and Risk Management Fu-Ming Lee, Ph.D. Director of Technology, GTC Technology Corporation, Houston, Texas: Recent Development of Extractive Distillation: A Distillation Alternative Jim Makris Director, Chemical Emergency Preparedness and Prevention Office, U.S. Environmental Protection Agency, Washington, D.C.: Process Safety and Risk Management Regulations: Impact on Process Safety M. Sam Mannan, Ph.D., P.E. Associate Professor of Chemical Engineering, Mary Kay O’Connor Process Safety Center, Texas A&M University, College Station, Texas: Fundamentals of Process Safety and Risk Management; Process Safety and Risk Management Regulations: Impact on Process Industry Rajen M. Patel, Ph.D. Technical Leader, The Dow Chemical Company, Polyolefins Research, Freeport, Texas: Structures, Properties, and Applications of Polyolefins Produced by Single-Site Catalyst Technology H. James Overman Dow Chemical Company, Freeport, Texas: Process Safety and Risk Management Regulations: Impact on Process Industry iii

iv

Contributors to Volume 69 Michael V. Pishko, Ph.D. Assistant Professor, Department of Chemical Engineering, Texas A&M University, College Station, Texas: Recent Advances in Biomaterials Alan W. Weimer, Ph.D., P.E. Professor of Chemical Engineering, University of Colorado, Boulder, Colorado: Effect of Pressure and Temperature in Bubbling Fluidized Beds

CONTENTS OF VOLUME 69 Contributors to Volume 69

iii

Conversion to SI Units

vii

Bringing Costs up to Date

ix

Circulating Fluidized Bed Reactors: Basic Concepts and Hydrodynamics Ray A. Cocco

1

Effect of Pressure and Temperature in Bubbling Fluidized Beds Alan W. Weimer

35

Fundamentals of Process Safety and Risk Management M. Sam Mannan, Dennis Hendershot, and Trevor A. Kletz

49

Introduction to the Selective Catalytic Reduction Technology Manfred Grove

94

Nanophase Materials in Chemical Process Avery N. Goldstein and David M. Fishbach

150

Process Safety and Risk Management Regulations: Impact on Process Industry M. Sam Mannan, Jim Makris, and H. James Overman

168

Recent Advances in Biomaterials Michael V. Pishko

194

Recent Development of Extractive Distillation: A Distillation Alternative Fu-Ming Lee

207

Structure, Properties, and Applications of Polyolefins Produced by Single-Site Catalyst Technology Rajen M. Patel and Steve Chum

231

v


To convert from

To

Multiply by

acre angstrom are atmosphere bar barrel (42 gallon) Btu (International Steam Table) Btu (mean) Btu (thermochemical) bushel calorie (International Steam Table) calorie (mean) calorie (thermochemical) centimeter of mercury centimeter of water cubit degree (angle) denier (international) dram (avoirdupois) dram (troy) dram (U.S. fluid) dyne electron volt erg fluid ounce (U.S.) foot furlong gallon (U.S. dry) gallon (U.S. liquid) gill (U.S.) grain gram horsepower horsepower (boiler) horsepower (electric) hundred weight (long) hundred weight (short) inch inch mercury inch water kilogram force kip knot (international) league (British nautical) league (statute)

square meter (m 2 ) meter (m) square meter (m 2 ) newton/square meter (N/m 2 ) newton/square meter (N/m 2 ) cubic meter (m 3 ) joule (J) joule (J) joule (J) cubic meter (m 3 ) joule (J) joule (J) joule (J) newton/square meter (N/m 2 ) newton/square meter (N/m 2 ) meter (m) radian (rad) kilogram/meter (kg/m) kilogram (kg) kilogram (kg) cubic meter (m 3 ) newton (N) joule (J) joule (J) cubic meter (m 3 ) meter (m) meter (m) cubic meter (m 3 ) cubic meter (m 3 ) cubic meter (m 3 ) kilogram (kg) kilogram (kg) watt (W) watt (W) watt (W) kilogram (kg) kilogram (kg) meter (m) newton/square meter (N/m 2 ) newton/square meter (N/m 2 ) newton (N) newton (N) meter/second (m/s) meter (M) meter (m)

4.046 ⫻ 10 3 1.0 ⫻ 10⫺10 1.0 ⫻ 10 2 1.013 ⫻ 10 5 1.0 ⫻ 10 5 0.159 1.055 ⫻ 10 3 1.056 ⫻ 10 3 1.054 ⫻ 10 3 3.52 ⫻ 10⫺2 4.187 4.190 4.184 1.333 ⫻ 10 3 98.06 0.457 1.745 ⫻ 10⫺2 1.0 ⫻ 10⫺7 1.772 ⫻ 10⫺3 3.888 ⫻ 10⫺3 3.697 ⫻ 10⫺6 1.0 ⫻ 10⫺5 1.60 ⫻ 10⫺19 1.0 ⫻ 10⫺7 2.96 ⫻ 10⫺5 0.305 2.01 ⫻ 10 2 4.404 ⫻ 10⫺3 3.785 ⫻ 10⫺3 1.183 ⫻ 10⫺4 6.48 ⫻ 10⫺5 1.0 ⫻ 10⫺3 7.457 ⫻ 10 2 9.81 ⫻ 10 3 7.46 ⫻ 10 2 50.80 45.36 2.54 ⫻ 10⫺2 3.386 ⫻ 10 3 2.49 ⫻ 10 2 9.806 4.45 ⫻ 10 3 0.5144 5.559 ⫻ 10 3 4.83 ⫻ 10 3

vii

viii


To convert from

To

Multiply by

light year liter micron mil mile (U.S. nautical) mile (U.S. statute) millibar millimeter mercury oersted ounce force (avoirdupois) ounce mass (avoirdupois) ounce mass (troy) ounce (U.S. fluid) pascal peck (U.S.) pennyweight pint (U.S. dry) pint (U.S. liquid) poise pound force (avoirdupois) pound mass (avoirdupois) pound mass (troy) poundal quart (U.S. dry) quart (U.S. liquid) rod roentgen second (angle) section slug span stoke ton (long) ton (metric) ton (short, 2000 pounds) torr yard

meter (m) cubic meter (m 3 ) meter (m) meter (m) meter (m) meter (m) newton/square meter (N/m 2 ) newton/square meter (N/m 2 ) ampere/meter (A/m) newton (N) kilogram (kg) kilogram (kg) cubic meter (m 3 ) newton/square meter (N/m 2 ) cubic meter (m 3 ) kilogram (kg) cubic meter (M 3 ) cubic meter (m 3 ) newton second/square meter (N ⋅ s/m 2 ) newton (N) kilogram (kg) kilogram (kg) newton (N) cubic meter (m 3 ) cubic meter (m 3 ) meter (m) coulomb/kilogram (c/kg) radian (rad) square meter (m 2 ) kilogram (kg) meter (m) square meter/second (m 2 /s) kilogram (kg) kilogram (kg) kilogram (kg) newton/square meter (N/m 2 ) meter (m)

9.46 ⫻ 10 15 0.001 1.0 ⫻ 10⫺6 2.54 ⫻ 10⫺6 1.852 ⫻ 10 3 1.609 ⫻ 10 3 100.0 1.333 ⫻ 10 2 79.58 0.278 2.835 ⫻ 10⫺2 3.11 ⫻ 10⫺2 2.96 ⫻ 10⫺5 1.0 8.81 ⫻ 10⫺3 1.555 ⫻ 10⫺3 5.506 ⫻ 10⫺4 4.732 ⫻ 10⫺4 0.10 4.448 0.4536 0.373 0.138 1.10 ⫻ 10⫺3 9.46 ⫻ 10⫺4 5.03 2.579 ⫻ 10⫺4 4.85 ⫻ 10⫺6 2.59 ⫻ 10 6 14.59 0.229 1.0 ⫻ 10⫺4 1.016 ⫻ 10 3 1.0 ⫻ 10 3 9.072 ⫻ 10 2 1.333 ⫻ 10 2 0.914

Bringing Costs up to Date Cost escalation via inflation bears critically on estimates of plant costs. Historical costs of process plants are updated by means of an escalation factor. Several published cost indexes are widely used in the chemical process industries: Nelson Cost Indexes (Oil and Gas J.), quarterly Marshall and Swift (M&S) Equipment Cost Index, updated monthly CE Plant Cost Index (Chemical Engineering), updated monthly ENR Construction Cost Index (Engineering News-Record), updated weekly Vatavuk Air Pollution Control Cost Indexes (VAPCCI) (Chemical Engineering), updated quarterly All of these indexes were developed with various elements such as material availability and labor productivity taken into account. However, the proportion allotted to each element differs with each index. The differences in overall results of each index are due to uneven price changes for each element. In other words, TABLE 1 Chemical Engineering and Marshall and Swift Plant and Equipment Cost Indexes since 1950 Year

CE Index

M&S Index

Year

CE Index

M&S Index

1950 1951 1952 1953 1954 1955 1956 1957 1958 1959 1960 1961 1962 1963 1964 1965 1966 1967 1968 1969 1970 1971 1972

73.9 80.4 81.3 84.7 86.1 88.3 93.9 98.5 99.7 101.8 102.0 101.5 102.0 102.4 103.3 104.2 107.2 109.7 113.6 119.0 125.7 132.3 137.2

167.9 180.3 180.5 182.5 184.6 190.6 208.8 225.1 229.2 234.5 237.7 237.2 238.5 239.2 241.8 244.9 252.5 262.9 273.1 285.0 303.3 321.3 332.0

1973 1974 1975 1976 1977 1978 1979 1980 1981 1982 1983 1984 1985 1986 1987 1988 1989 1990 1991 1992 1993 1994 1995 1996 1997

144.1 165.4 182.4 192.1 204.1 218.8 238.7 261.2 297.0 314.0 316.9 322.7 325.3 318.4 323.8 342.5 355.4 357.6 361.3 358.2 359.2 368.1 381.1 381.7 386.5

344.1 398.4 444.3 472.1 505.4 545.3 599.4 659.6 721.3 745.6 760.8 780.4 789.6 797.6 813.6 852.0 895.1 915.1 930.6 943.1 964.2 993.4 1027.5 1039.2 1056.8

ix

x

Bringing Costs up to Date TABLE 2 Nelson-Farrar Inflation Petroleum Refinery Construction Indexes since 1946 (1946 ⫽ 100)

Date

Materials Component

Labor Component

Miscellaneous Equipment

Nelson Inflation Index

1946 1947 1948 1949 1950 1951 1952 1953 1954 1955 1956 1957 1958 1959 1960 1961 1962 1963 1964 1965 1966 1967 1968 1969 1970 1971 1972 1973 1974 1975 1976 1977 1978 1979 1980 1981 1982 1983 1984 1985 1986 1987 1988 1989 1990 1991

100.0 122.4 139.5 143.6 149.5 164.0 164.3 172.4 174.6 176.1 190.4 201.9 204.1 207.8 207.6 207.7 205.9 206.3 209.6 212.0 216.2 219.7 224.1 234.9 250.5 265.2 277.8 292.3 373.3 421.0 445.2 471.3 516.7 573.1 629.2 693.2 707.6 712.4 735.3 739.6 730.0 748.9 802.8 829.2 832.8 832.3

100.0 113.5 128.0 137.1 144.0 152.5 163.1 174.2 183.3 189.6 198.2 208.6 220.4 231.6 241.9 249.4 258.8 268.4 280.5 294.4 310.9 331.3 357.4 391.8 441.1 499.9 545.6 585.2 623.6 678.5 729.4 774.1 824.1 879.0 951.9 1044.2 1154.2 1234.8 1278.1 1297.6 1330.0 1370.0 1405.6 1440.4 1487.7 1533.3

100.0 114.2 122.1 121.6 126.2 145.0 153.1 158.8 160.7 161.5 180.5 192.1 192.4 196.1 200.0 199.5 198.8 201.4 206.8 211.6 220.9 226.1 228.8 239.3 254.3 268.7 278.0 291.4 361.8 415.9 423.8 438.2 474.1 515.4 578.1 647.9 622.8 656.8 665.6 673.4 684.4 703.1 732.5 769.9 797.5 827.5

100.0 117.0 132.5 139.7 146.2 157.2 163.6 173.5 179.8 184.2 195.3 205.9 213.9 222.1 228.1 232.7 237.6 243.6 252.1 261.4 273.0 286.7 304.1 329.0 364.9 406.0 438.5 468.0 522.7 575.5 615.7 653.0 701.1 756.6 822.8 903.8 976.9 1025.8 1061.0 1074.4 1089.9 1121.5 1164.5 1195.9 1225.7 1252.9

Bringing Costs up to Date

xi

TABLE 2 Continued

Date

Materials Component

Labor Component

Miscellaneous Equipment

Nelson Inflation Index

1992 1993 1994 1995 1996 1997

824.6 846.7 877.2 918.0 917.1 923.9

1579.2 1620.2 1664.7 1708.1 1753.5 1799.5

837.6 842.8 851.1 879.5 903.5 910.5

1277.3 1310.8 1349.7 1392.1 1418.9 1449.2

the total escalation derived by each index will vary because different bases are used. The engineer should become familiar with each index and its limitations before using it. Table 1 compares the CE Plant Index with the M&S Equipment Cost Index. Table 2 shows the Nelson-Farrar Inflation Petroleum Refinery Construction Indexes since 1946. It is recommended that the CE Index be used for updating total plant costs and the M&S Index or Nelson-Farrar Index for updating equipment costs. The Nelson-Farrar Indexes are better suited for petroleum refinery materials, labor, equipment, and general refinery inflation. Since C B ⫽ C A (B/A) n

(1)

Here, A ⫽ the size of units for which the cost is known, expressed in terms of capacity, throughput, or volume; B ⫽ the size of unit for which a cost is required, expressed in the units of A; n ⫽ 0.6 (i.e., the six-tenths exponent); C A ⫽ actual cost of unit A; and C B ⫽ the cost of B being sought for the same time period as cost C A . To approximate a current cost, multiply the old cost by the ratio of the current index value to the index at the date of the old cost: C B ⫽ C A I B /I A

(2)

Here, C A ⫽ old cost; I B ⫽ current index value; and I A ⫽ index value at the date of old cost. Combining Eqs. (1) and (2), C B ⫽ C A (B/A) n (I B /I A )

(3)

For example, if the total investment cost of plant A was $25,000,000 for 200million-lb/yr capacity in 1974, find the cost of plant B at a throughput of 300 million lb/yr on the same basis for 1986. Let the sizing exponent, n, be equal to 0.6. From Table 1, the CE Index for 1986 was 318.4, and for 1974 it was 165.4. Via Eq. (3),

xii

Bringing Costs up to Date TABLE 3 Vatavuk Air Pollution Control Cost Indexes (VAPCCI). First Quarter 1994 ⫽ 100.0 (index values have been rounded to the nearest tenth). Control Device

1994 (Avg.)

1995 (Avg.)

1996 (Avg.)

Carbon adsorbers Catalytic incinerators Electrostatic precipitators Fabric filters Flares Gas absorbers Mechanical collectors Refrigeration systems Regenerative thermal oxidizers Thermal incinerators Wet scrubbers

101.2 102.0 102.8 100.5 100.5 100.8 100.3 100.5 101.4 101.3 101.3

110.7 107.1 108.2 102.7 107.5 105.6 103.0 103.0 104.4 105.9 112.5

106.4 107.0 108.0 104.5 104.9 107.8 103.3 104.4 106.3 108.2 119.8

C B ⫽ C A (B/A) n (I B /I A ) ⫽ 25.0 (300/200) 0.6 (318.4/165.4) ⫽ $61,200,000 Table 3 shows the Vatavuk Air Pollution Control Cost Indexes (VAPCCI) since 1994. For details, see the Vatavuk Air Pollution Control Cost Indexes article in volume 61. Editor’s note: For a more thorough explanation of updating costs, see the article, ‘‘Tower Cost Updating’’ in volume 58.

john j. mcketta

Encyclopedia of Chemical Processing and Design 69

Circulating Fluidized Bed Reactors: Basic Concepts and Hydrodynamics

Introduction Circulating fluidized beds (CFBs) consist of two basic designs, as shown in Fig. 1. One design involves a fast-fluidized bed where high gas velocities convey a substantial amount of solids to one or more cyclones. The separated particles are fed back to the fluidized bed using a standpipe. The second basic design uses a riser to convey solids to one or more cyclones. The separated particles are fed to an optional fluidized bed and then back to the riser. Solids flow rates can be controlled using nonmechanical L- and J-valves or using a mechanical slide valve. The large-scale commercial realization of CFBs occurred in the early 1940s, although some coal gasification was done in a fluidized bed as early as 1926 [1]. With the increased demand for gasoline during World War II, major efforts were underway to develop reactors to crack petroleum feedstocks into usable fuels more productively than the moving bed or snake reactors (i.e., the Houndry Process) used at that time. The result was a fluidized catalyst cracker (FCC), where high catalyst circulation rates allowed a balance between the exothermic burning of coke on the catalyst in the regenerator and the endothermic hydrocracking of petroleum in the reactor. The continuous circulation or regeneration of catalyst provided fresh catalyst for petroleum cracking and thereby resulted in high sustainable productivities. With the addition of a stripping section after the reactor, even higher yields were obtained. The addition of steam, CO 2, or other inerts would remove the product from and around catalyst particles flowing toward the regenerator. Today, the evolution of the FCC unit has results into several basic designs, as shown in Fig. 2. In 1960, circulating fluidized beds contributed to another breakthrough process for the petroleum and chemical industry. Standard Oil of Ohio (SOHIO) developed a fluidized-bed reactor for the ammoxidation of propene to acrylonitrile. Previous technology was done in tube-and-shell fixed-bed reactors. However, the high heat of reaction of 160 kcal/mol limited the economic feasibility of those units. The high heat transfer characteristic of fluidized-bed reactors made them ideal for the production of acrylonitrile. Today, nearly all large-scale acrylonitrile plants are based on the SOHIO design, with capacities up to 180,000 tons per year [4]. The greatest challenge in developing the SOHIO process was in the management of backmixing. The inherent hydrodynamics of fluidized beds, where solids and, to a lesser extent, gas circulate from the top of the bed to the bottom, then to the top again, would have a deleterious effect on acrylonitrile selectivity. To overcome backmixing, SOHIO developed sieve trays to compartmentalize the gas flow in the fluidized-bed reactor to resemble a more plug-flow characteristic [5]. In 1979, SOHIO redesigned the acrylonitrile reactor to more of a ‘‘tube-and-shell’’ fluidized-bed unit [6], as shown in Fig. 3. 1

2

Circulating Fluidized Bed Reactors

FIG. 1 Basic design of circulating fluidized beds.

FIG. 2 Typical FCC units based on the designs of (a) Standard Oil Development, (b) UOP, (c) Kellogg, and (d) Exxon. (Adapted from Refs. 2 and 3.)


3

FIG. 3 Two-dimensional schematic of the SOHIO acrylonitrile processes. (Adapted from Ref. 5.)

During the late 1970s and early 1980s oil crisis, circulating fluidized beds found applications in coal combustion. The high-heat-transfer capabilities of these reactors resulted in lower operating temperatures, thereby reducing NO x and SO 2 emissions. In addition, the high gas velocities resulted in significant turbulence, which provided uniform temperatures in the combustor. With the surplus of oil starting in the late 1980s, fluidized-bed combustors became economically less attractive. As of the early 1990s, only Dynergy (via the Destec process) and Lurgi and Ahlstrom are practicing this technology [7]. Today, circulating fluidized beds are used in a wide array of chemical processes, as shown in Table 1. With fluidized beds having the unique distinction of excellent heat transfer and continuous in situ regeneration, the economic attractiveness of processing thermally sensitive chemicals or using catalysts that require TABLE 1 Some Fluidized and Circulating Fluidized Bed Reactor Processes Product Acrolynitrile Aniline Chloromethanes Goal gasification Hydrocyanic acid Maleic anhydride Maleic anhydride Maleic anhydride Perchlorethlyene Phthalic anhydride Polyethylene Synthesis gas Vinyl chloride

Process

Developer

Propene ammoxidation Nitrobenzene hydrogenation Cat. oxidaton of methane Oxidation Ammono-dehydrogenation Butane oxidation Butane oxidation Butene oxidation Chlorination Naphthalene oxidation Ethlyene polymerization Fisher–Tropsch Ethylene chlorination, oxychlorination

SOHIO BASF, Cyanamid, Lonza Asahi Glass Winkler Shawinigan Alusuisse & Lummus (Alma Process) DuPont Mitsubishi Diamond Shamrock Badger/Sherwin-Williams Union Carbide (Unipol) SASOL, Kellogg Ethyl, Hoechst, Mitsui, Toatsu, Monsanto

4

Circulating Fluidized Bed Reactors frequent regeneration are more realized. Once the obstacles of backmixing, mass transfer, and attrition have been addressed, these reactors often set the standards in reactor design.

Basic Concepts As the gas velocity through a bed of solids increases, the bed undergoes several regimes, as shown in Fig. 4. At first, the gas velocity is insufficient to fluidize the particles and the bed remains fixed. With increasing velocity and under ideal conditions, the fixed bed expands smoothly and uniformly. Particles move in a limited fluidlike fashion and the bed pressure drop becomes constant. At this point, the bed is commonly referred to as undergoing minimum fluidization. Further increases in gas velocity results in further bed expansion and particles appear to freely move throughout the bed. The gas permeates through the bed without the formation of bubbles. This regime is referred to as smooth fluidization and is only observed for Geldart Group A powders (see Appendix A). These powders require noticeably higher gas velocities to promote the formation of gas bubbles after minimum fluidization. In contrast, Group B powders begin bubbling shortly after minimum fluidization. Group C powders, being cohesive, may even show signs of bubbling prior to minimum fluidization; however, this is usually the result of channeling. The onset of bubbles in the fluidized bed is commonly referred to as bubbling fluidization. Here, gas bubbles form at or near the distributor and grow to a maximum bubble size as they propagate through the bed. The top of the fluidized bed is still well defined, as it was in the minimum and smooth fluidization regimes. The pressure drop across the bed is still constant, on average, but starts exhibiting large, but regular, fluctuations with time. As the gas velocity continues to increase, the top of the bed becomes less defined. Large amounts of particles are ejected into the freeboard region above the bed. Concurrently, sizeable regions of voidage and particle clusters are seen

FIG. 4 Various fluidization regimes with increasing superficial gas velocity.


5

in the bed itself. For Group A and B powders, this transition from the bubbling fluidized-bed regime is called the onset of turbulent fluidization. Group C and D powders may show a slugging behavior prior to the turbulent fluidization regime. During fast fluidization, the gas velocity is sufficient enough that the surface of the bed can no longer be discerned. Particle density is still higher at the bottom of the unit compared to the top, suggesting that some sort of bed exists. Particle clusters and streamers are readily observed and, in some cases, a core–annulus radial variation in particle density begins to take shape. Particle entrainment is high and the total disengagement height may be well beyond the physical dimensions of the fluidized-bed unit. To overcome the losses of particles due to entrainment, cyclones may be used to capture entrained particles and recirculate them back into the bed. At very high gas velocities, nearly all the particles are entrained from the bed. This regime is commonly referred to as pneumatic conveying. In this regime, axial variation in particle density is no longer observed, except maybe in entrance and exit regions. Radial variation in particle density can vary dramatically and range from a core–annulus profile to a uniform profile. For dense systems, clusters and streamers are readily observed. Thus, for gas–solid systems, increases in the gas velocity results in dramatic and sometimes sharp transitions in the hydrodynamics. In the design of fluidized beds, it is crucial that one knows the fluidized regime that will exist at operating conditions. The simple transition from one regime to another can have significant impacts on reaction, heat transfer, attrition, and entrainment rates. For circulating fluidized beds, where several regimes may exist in a single unit (i.e., from conveying in the riser to a bubbling fluidized-bed regime in the regenerator), knowledge of the fluidization regimes is paramount. In order to gain better understanding of these regimes, the methodology used to determine the onset of each fluidization regime is discussed in the following sections. Keep in mind that most of the correlations are empirical and may not fully represent every system. With the cost of these units running in the tens of millions of dollars for large-scale plants, experimental validation of the expected regimes is critical when designing these processes.

Minimum and Smooth Fluidization As a gas permeates through a fixed or packed bed, the pressure drop can be described by the Ergun equation [8]: (1 ⫺ ε bp )2µu g (1 ⫺ ε bp )ρ g u 2g ∆P g c ⫽ 150 ⫹ 1.75 L ε 3bp (Φd p ) 2 ε 3bp (Φd p ) 2

(1)

With increasing gas velocity, the bed reaches a point where the drag force exceeds the force of gravity on the particles. The particles become mobile and the bed becomes fluidized. The gas velocity at the onset of this type of fluidization is referred to as the minimum fluidization velocity or u mf. Assuming that the weight

6

Circulating Fluidized Bed Reactors of the particles in the fluidized bed corresponds to ∆P/L, the Ergun equation can be written as

冢

冣

1.75 d p u mf ρ g ε 3mf Φ µ

2

⫹

冢

冣

150(1 ⫺ ε mf ) d p u mf ρ g d 3 gρ (ρ ⫺ ρ g ) ⫽ p g 2s 3 2 ε mf Φ µ µ

(2)

or 150(1 ⫺ ε mf ) 1.75 2 Re p,mf ⫹ Re p,mf ⫽ Ar ε 3mf Φ ε 3mf Φ 2

(3)

where the Archimedes number, Ar, is defined as

Ar ⫽

d 3pgρ g (ρ s ⫺ ρ g ) µ2

(4)

and Re p is the particle Reynolds number having the expression

Re p,mf ⫽

d p u mf ρ g µ

(5)

Equation (3) can be written as a quadratic with the coefficients K 1 and K 2 having the form K 1 Re 2p,mf ⫹ K 2 Re p, mf ⫽ Ar

(6)

By solving for Re p, mf, Eq. (6) can be rewritten as

Re p, mf ⫽

冤冢冣冢冣冥 ⫺ 冢2KK 冣 K1 2K 2

2

⫹

Ar K1

0.5

1

(7)

2

where the particle Reynolds number at minimum fluidization is a simple function of the Archimedes number and two constants (K 1 /2K 2 and 1/K 1 ). Many correlations for the minimum fluidization velocity are based on Eq. (7)


7

TABLE 2 K1 and K2 Values for Eq. (7) Reference Wen and Yu [9] Richardson [10] Saxena and Vogel [11] Bubu et al. [12] Grace [13] Chitester et al. [14]

K2 /2K1

1/K1

33.7 25.7 25.3 25.3 27.2 28.7

0.0408 0.0365 0.0571 0.0651 0.0408 0.0494

Comments For fine particles Dolomite at high temperature and pressure

For large particles

Source: Adapted from Ref. 7.

for the constants K 1 /2K 2 and 1/K 1. These constants are presented in Table 2 for a wide range of studies. For typical Geldart Group A powder, the constants of Wen and Yu are most often used. However, these correlations are specific to a group of particles with common characteristics and may not represent a less-thanideal particle morphology and texture. The minimum fluidization velocity can be experimentally determined by measuring the pressure drop across a bed of particles with increasing superficial gas velocity. For smooth, round, and noncohesive particles, the pressure drop increases linearly with gas velocity until the minimum fluidization velocity is reached. With further increases in the gas velocity, the pressure drop remains constant. Hence, the minimum fluidization velocity is the intersection of the linearly increasing line with the constant-pressure-drop line. Figure 5–8 demonstrate the results of such an experiment. Figure 5 is the pressure-drop curve for alumina particles with a mean particle diameter of 60 µm in a 4.5-in.-inner diameter fluidized bed unit. The minimum fluidization velocity for these particles was determined to be 6.5 cm/min. When measuring the minimum fluidization velocity, less scatter in the data is obtained from larger or higher beds. The scatter in Fig. 5 suggests that perhaps a higher bed should have been used. The diameter of the fluidized bed used in this type of experiment is also

FIG. 5 Minimum fluidization curve for smooth and round alumina particles, d p,ave ⫽ 60 µm.

8


FIG. 6 Minimum fluidization curve for smooth and round alumina particles, d p,ave ⫽ 60 µm, with increasing and decreasing gas velocities.

critical to obtaining accurate data. For Geldart Group A powders, bed diameters less than 3 in. can result in experimental data that are influenced by frictional effects at the wall. For the coarser Group B powders, the minimum diameter is much larger. Figure 6 shows two pressure drop versus superficial gas velocity curves for the same particles used in Fig. 5. The black data points are the pressure drop with increasing gas velocity and the gray data points are the subsequent pressure-drop measurements with decreasing gas velocities. For round, smooth, and noncohesive particles, the two curves should overlap each other, as shown in Fig. 6. However, for irregular, rough, or cohesive particles, a hysterisis effect is typically observed. This is obvious in Fig. 7 for rough and irregularly shaped alumina particles with

FIG. 7 Minimum fluidization curve for rough alumina particles, d p,ave ⫽ 92 µm.


9

FIG. 8 Minimum fluidization curve for cohesive iron catalyst particles, d p,ave ⫽ 68 µm.

a mean particle diameter of 92 µm. This resulted in higher solid shear forces during fluidization such that the pressure drop is dependent on previous conditions or is path dependent. Figure 8 shows the pressure-drop curve for a catalyst supported on alumina with a mean particle diameter of 68 µm. Here, it is almost impossible to detect the minimum fluidization velocity. High cohesive forces result in a fluidized bed prone to channeling. Each peak or spike in Fig. 8 is the result of another channel achieving fluidization while the majority of the bed remains fixed. This behavior is typically of Group C powders.

Bubbling Fluidization As discussed earlier, beds with Group A powders pass from minimum fluidization to smooth fluidization to bubbling fluidization with increasing gas velocity. Group B and D powders exhibit bubbling fluidization at the onset of minimum fluidization. Oddly enough, gas bubbles in all fluidized beds behave similarly to gas bubbles in low-viscosity liquids [7]. Large gas bubbles are typically spherical on top and flattened or even inverted on the bottom; smaller bubbles tend to be completely spherical. As in liquid systems, gas bubbles in fluidized beds can coalesce into larger bubbles or split into smaller bubbles, depending on bed conditions. Also, as gas bubbles approach the top of a fluidized bed, they collapse such that solids are propagated into the freeboard region. Higher pressures or temperatures result in a decrease in the maximum bubble size (due to changes in the gas physical properties) and tend to make fluidization smoother [1]. The minimum velocity for bubble formation is referred to as the minimum bubbling velocity or U mb. For Geldart Group A and C powders, Abrahamsen and Geldart [15] proposed that the minimum bubbling velocity can be calculated from

10

Circulating Fluidized Bed Reactors 0.52 u mb 2300ρ 0.13 g ug ⫽ 0.8 exp(0.72P 45 µm ) in SI units u mf d p,ave (ρ s ⫺ ρ s )0.93

(8)

This expression is based on observations that the minimum bubble velocity is strongly dependent on P 45 µm, the probability of finding a particle with a diameter less than 45 µm. For calculations where less information regarding the system is known, Eq. (8) can be simplified into the form u mb

冢冣

ρ ⫽ 33d p g µg

0.1

in SI units

(9)

For Geldart Group B and D powders, the minimum bubbling velocity is near the minimum fluidization velocity. Thus, the onset of fluidization and the formation of gas bubbles nearly coincide and u mb ⫽ u mf

(10)

As gas bubbles rise through the fluidized bed, the bubble size increases until a maximum or equilibrium size is achieved, provided that the bed is high enough. For Group A and B powders, Davidson and Harrison [16] proposed that the maximum stable bubble size can be determined from d b,max ⫽

2u t g

(11)

where u t is the terminal free-fall velocity of the particle (see Appendix B). Geldart [1] found that Eq. (11) provided a better fit to experimental data if an effective diameter, d′p, was used to calculate the terminal velocity [i.e., u t ⫽ f(d′p )]. The effective diameter is defined as d′p ⫽ 2.7d p

(12)

For Group D powders, the maximum stable bubble size is so large that in realistic fluidized-bed applications the bubbles size is restricted by the bed diameter.

Turbulent Fluidization Further increases in the gas velocity result in a slugging or turbulent fluidized bed. If the bed diameter is small and the bed is sufficiently high, the fluidized bed will slug before entering the turbulent fluidization regime. For Group A, B, and D powders, slugging is basically the result of a bubble diameter that exceeds about two-thirds the bed diameter. The wall stabilizes the bubble such that almost the entire bed is translated up to the top of the bed or even higher. Group C powders may also exhibit slugging behavior even in large-diameter beds due to the cohesive forces. Thus, the larger and more cohesive the particles or the smaller the bed diameter, the higher the probability of a bed exhibiting slugging. For these cases,


11

the minimum slugging velocity, U ms, can be estimated using the expression of Stewart and Davidson [17], where u ms ⫽ u mf ⫹ 0.07(gD)0.5

(13)

For Group A or B powders in large-diameter beds, where the maximum bubble diameter is smaller than two-thirds the bed diameter, higher gas velocities result in the onset of turbulence fluidization directly from the bubbling bed regime. The onset of turbulent fluidization is defined as the point where the standard deviation of the pressure fluctuations reaches a maximum [18]. Grace and Bi [19] noted that fluidization at this point reflects the balance between bubble coalescence and breakup. Higher velocities cause this balance to shift toward a dominating bubble breakup mechanism. The minimum velocity for the onset of turbulence fluidization, u mt, can be calculated using the expression of Bi and Grace [20]:

u mt ⫽ 1.24 Ar 0.45

冢冣 µ ρg dp

(14)

where 2 ⬍ Ar ⬍ 1 ⫻ 108 or, in terms of dimensionless numbers, Re mt ⫽ 1.24 Ar 0.45

(15)

where 2 ⬍ Ar ⬍ 1 ⫻ 108.

Fast-Fluidization Regime Detecting the point of transition between the turbulent fluidized bed and the fastfluidized bed regimes has been under debate for many years. Yerushalmi [21] proposed that this transition occurs when a significant number of particles become entrained from the fluidized bed. This transition can be observed by detecting a sudden change of pressure drop with respect to entrainment rate for increasing superficial velocities. The superficial gas velocity corresponding to this point is referred to as the minimum transport velocity or u mr. Schnitzlein and Weinstein [22], however, were unable to determine u mr using Yerushalmi’s method. Their observations suggested that the value of u mr was strongly dependent on the location and the distance separating the two pressure taps used to measure the pressure drop. Bi and Grace [20] measured the entrainment rate versus superficial velocity for a wide range of fluidized-bed systems. They noted that the onset of fast fluidization correlated to the point where significant particle entrainment was observed. For Groups A and B powders, this minimum transport velocity can be estimated with the expression u mr ⫽ 1.53 Ar 0.45

冢冣 µ ρgdp

(16)

12

Circulating Fluidized Bed Reactors where 2 ⬍ Ar ⬍ 4 ⫻ 106. For the larger Group D powders, Eq. 16 may underpredict u mr compared to using the terminal velocity as the minimum transport velocity. Under these conditions, the minimum transport velocity should be set to equal the terminal velocity [19].

Pneumatic Conveying As the velocity continues to increase in a fast-fluidized bed, the axial transition between the dense and lean regions disappears. This transition marks the onset of the pneumatic conveying regime and the superficial velocity corresponding to this point is called the minimum conveying velocity or u mc. The onset of pneumatic conveying is readily measured by starting with a high superficial gas velocity and decreasing its value while holding the entrainment rate constant. The superficial gas velocity where the suspension collapse (i.e., choking) is observed corresponds to the minimum conveying velocity [19]. Yang [23] proposed that this velocity can be estimated with the equation

冢冣

2gD(ε ⫺4.7 ρ mc ⫺ 1) ⫽ 681,000 g u mc /ε mc ⫺ u t ρs

2.2

(17)

Regime Map The bed hydrodynamics associated with each powder classification used in fluidized beds are summarized in Fig. 9. Increasing the gas velocity through a powder bed results in the transition of several regimes, which are dependent on the particle properties. For Group A powders, increasing the gas velocity results in bed hydrodynamics that start in the fixed-bed regime and continue through smooth fluidization, bubbling fluidization, turbulent fluidization, fast fluidization, and, finally, pneumatic transport. Group B powders exhibit the same types of transition, with the exception that smooth fluidization is typically not observed; the onset of fluid-

FIG. 9 Flow regime map for various powders. Slugging for Group A and B powders depend on vessel diameter. Group C powders tend to slug and Group D powders almost always exhibit slugging.


13

ization corresponds with the onset of bubble formation. For Group C powders, smooth and bubbling fluidization are often replaced by channeling due to large solid stresses or large cohesive forces on these particles. Group C powders also have a higher tendency to exhibit slugging prior to turbulent fluidization. Group D powders often exhibit slugging during the onset of fluidization. The large particle sizes and/or high densities of Group D powder result in the formation of large maximum bubble diameters that often exceed the diameters of commercial-scale fluidized beds. In contrast, slugging for Group A, B, and, to a lesser extent, C powders may not be observed in large commercial units where the effective bed diameter often exceeds the maximum stable bubble size. In an effort to develop a design guide for fluidized unit operations, various attempts have been made to quantify the flow regimes for gas–solid flow. Reh [24] first made the attempt for a unified flow regime map by comparing the particle’s Reynolds number, Re p, to the inverse of the drag coefficient, 1/CD. Further modification to Reh’s map were made by Werther [25]. Li and Kwauk [26] and Avidan and Yerushalmi [27] took a different approach to developing a unified flow regime map by comparing the relationship of the superficial gas velocity to solids voidage. This approach was further modified by Rhodes [28]. Yet, another approach was taken by Leung [29], Klinzing [30], and Yang [23], who compared the superficial gas velocity to the solid flux. In each of these cases, however, the transition from the fast-fluidized bed to a pneumatic transport regime was not well defined. Grace [31] resolved this problem by generating a unified flow regime map based on the dimensionless variables d*p and u* of Zenz and Othmer [32]. By comparing the dimensionless particle diameter, d *p ⫽ Ar 1/3, with the dimensionless gas velocity, u* ⫽ Re p /Ar 1/3, Grace was able to discern the transition to pneumatic conveying. Kunii and Levenspiel [7] further modified Grace’s work by including the observations of van Deemter [33], Horio et al. [34], and Catipovic et al. [35]. Figure 10 shows the result of this effort. It is interesting to note that Kunii and Levenspiel’s [7] demarcation from fast fluidization to pneumatic conveying is not identical to that first proposed by Grace. Kunii and Levenspiel describe this transition as not being well defined. With the aid of Fig. 10 and the relatively simple calculation of d*p and u*, the flow behavior of most gas–solid systems can be readily obtained. Keep in mind, however, that the boundary for each regime shown in Fig. 10 should not be considered as sharp transitions, but more of a gray area between two adjacent regimes.

Pressure Balance in a CFB In designing a circulating fluidized bed (CFB), special attention needs to be given to the pressure-loop calculations. This is especially true in the proper design of standpipes and mechanical or nonmechanical valves (i.e., L-valves, J-valves, slide valves). The bed height in a standpipe counterbalances the dynamic pressure drop occurring in the riser or fast-fluidized-bed section. Mechanical or nonmechanical valves provide better response for this counterbalancing. For instance, an increase in the gas flow rate in a riser would require a constriction in a slide valve or less aeration in a L- or J-valve to prevent backflow. This correction results in a higher standpipe bed height and pressure drop. The higher pressure in the standpipe coun-

14


FIG. 10 Flow regime diagram proposed by Kunii and Levenspeil. (From Ref. 7.)

terbalances the increased pressure in the riser. Counterbalancing can also happen without the aid of mechanical or nonmechanical valves, but the response time is slow and the probability for backflow is much greater. Figures 11–13 give qualitative examples of the pressure-loop behavior in the CFBs shown in Fig. 1. The pressure drop in the riser or fluidized-bed section must equal the pressure drop in the cyclone, fluidized bed, standpipe, and control valve sections. For example, in Fig. 11, the pressure loop is defined as [(P 4 ⫺ P 3 )] fluid bed ⫹ [(P 5 ⫺ P 4 )] freeboard ⫽ [(P 6 ⫺ P 7 ) ⫹ (P 5 ⫺ P 6 )] cyclone

(18)

⫹ [(P 8 ⫺ P 9 ) ⫹ (P 7 ⫺ P 8 )] standpipe or ∆P| fluid bed ⫹ ∆P| freeboard ⫽ ∆P| cyclone ⫹ ∆P| standpipe

(19)

Assuming that particle frictional (both particle–particle and particle–wall) and particle acceleration effects are negligible, the pressure drop in the fluidized bed can be approximated by ∆P| fluid bed ⫽ L bρ s g(1 ⫺ ε)

(20)

where L b is the height of the fluidized bed. Similarly, the freeboard can be described


15

FIG. 11

Pressure loop for a CFB of Design I in Fig. 1.

using ∆P| freeboard ⫽ ρ s g

冮

Lu

Lb

[1 ⫺ ε(z)] dz ⬇ ρ s g(1 ⫺ ε)(L u ⫺ L b )

(21)

where L u is the height of the unit. Because the voidage profile, ε(z), in the freeboard region is rarely known, Eq. (21) can be approximated by using the average voidage in the freeboard ε. For the right-hand side of Eq. (18), the pressure drop in the cyclone can be determined using the expression described by Muschelknautz and Greif [36], where ∆P| cyclone ⫽ ∆P| wall friction ⫹ ∆P| vortex losses

(22)

The first term of this expression corresponds to the friction losses at the wall due to gas. The second terms accounts for momentum losses in the vortex. Under the assumption that the flow over the wall of the cyclone is similar to flow over a flat plate, the pressure drop due to wall friction in a cyclone can be calculated using ∆P| wall friction ⫽ λ p

ρ g A surface (u 0 u b )0.67 in SI units 1.8Q g

(23)

The pressure drop in the vortex depends on the average velocity in the vortex tube and the tangential velocity in the vortex tube or exit. This pressure drop can be

16

Circulating Fluidized Bed Reactors calculated from

冤

∆P| vortex losses ⫽ ρ g v 20 2 ⫹ 3

冢冣冢冣冥 u0 v0

1.33

u0 v0

2

in SI units

(24)

where the average velocity in the vortex tube is v0 ⫽

V˙ g πr 20

(25)

For the standpipe, the pressure drop is determined in a method similar to that used for the fluidized bed under the same assumptions. The pressure drop in the lean phase of the standpipe can be estimated as ∆P| standpipe, lean phase ⫽ ρ s g

冮

Ls , 0

L s,b

[(1 ⫺ ε)(z)] dz ⬇ ρ s g(1 ⫺ ε)(L s , 0 ⫺ L s,b )

(26)

where L s,0 and L s,b are the height of the standpipe and the dense-phase bed in the standpipe, respectively. For the dense-phase region of the standpipe, the pressure drop is approximated as ∆P| standpipe, dense phase ⫽ ρ s g(1 ⫺ ε)L s,b For the CFB in Fig. 12, the pressure loop has the expression

FIG. 12

Pressure loop for CFB with a fluidized bed as shown in Design II of Fig. 1.

(27)


17

FIG. 13 Pressure loop for CFB with no fluidized bed as shown in Design III of Fig. 1.

(P 4 ⫺ P 3 ) riser ⫽ [(P 5 ⫺ P 4 ) ⫹ (P 6 ⫺ P 5 )] cyclone ⫹ [(P 7 ⫺ P 6 ) ⫹ (P 8 ⫺ P 7 )] 1st Standpipe (28) ⫹ [(P 9 ⫺ P 8 ) ] 2nd standpipe ⫹ [(P 10 ⫺ P 9 ) ⫺ (P 3 ⫺ P 10 )] L-valve or ∆P| riser ⫽ ∆P| cyclone ⫹ ∆P| 1st standpipe ⫹ ∆P| 2nd standpipe ⫹ ∆P| L-valve

(29)

Similarly, the pressure loop for the CFB in Fig. 13 can be described as (P 4 ⫺ P 3 ) riser ⫽ [(P 5 ⫺ P 4 ) ⫹ (P 6 ⫺ P 5 )] cyclone ⫹ [(P 7 ⫺ P 6 ) ⫹ (P 8 ⫺ P 7 )] standpipe ⫹ [(P 9 ⫺ P 8 ) ⫹ (P 3 ⫺ P 9 )] L-valve

(30)

∆P| riser ⫽ ∆P| cyclone ⫹ ∆P| standpipe ⫹ ∆P| L-valve

(31)

or

A similar set of expressions can be obtained for the CFB in Fig. 13. Compared to the pressure-loop calculations used for the system in Fig. 11, the only new expressions needed to complete the pressure-loop calculations in Figs. 12 and 13 are the pressure drop across the L-valve. Yang and Knowlton [37] noted that the

18

Circulating Fluidized Bed Reactors pressure drop across an L-valve is similar to that used by Jones and Davidson [38] for a mechanical valve where ∆P| mechanical

冢冣

1 Ws ⫽ 2ρ s (1 ⫺ ε mf ) C D A 0

2

(32)

For mechanical valves, the valve opening area, A o, may vary during operation but is always known. For an L-valve, however, the opening area, A o, depends on the amount of aeration and is not known. Yang and Knowlton [37] found an empirical expression for this opening area of the L-valve based on 158 experimental data points. Their results showed that the opening area can be expressed as A o ⫽ 1.41

Q aeration πD 2L-valve L L-valve ⫺ 0.0623 in Foot Pounds Seconds (FPS) units ut ut (33)

Thus, pressure drops through a CFB can be estimated and can provide the basic foundation for design. However, as these units become more complicated or are operated at higher and higher solid circulation rates, this procedure may not be enough. Indeed, the pressure-drop calculations presented assume that frictional and acceleration effects are negligible. Although this may be a safe assumption for rough estimates of the pressure loop, detailed design calculations cannot neglect these effects. In addition, particle properties such as roughness, sphericity, and particle size distribution can also have a significant effect on the pressure drop. Indeed, the addition of fines to a Group A powder can result in reduced pressure drops for fluidized beds and risers. Using computation fluid dynamics, Sinclair and Mallo [39] demonstrated that this may be due to the fact that smaller particles dampen the wakes generated by larger particles or bubbles. Hence, a clear understanding of the hydrodynamics is needed to fully describe pressure loss in CFBs. Fortunately, this has been the focus of intensive research in the last few years.

Gas–Solid Hydrodynamics In general, the hydrodynamics of a riser can be divided into macro-scale and mesoscale flow behavior. Macro-scale behavior is mostly concern with solids concentration profiles and solids velocity on a large scale. Risers typically exhibit a wide and diverse range of axial and radial solids profiles that are highly dependent on operating and design conditions. For instance, the design of the entrance and exit regions of a riser can influence the solid profile throughout the riser. Hence, it is important to understand this macro-scale behavior in order to provide the correct riser design for the desired solids concentration and velocity profiles. To make matters even more complicated, fast-fluidized beds and risers have shown evidence of dynamic meso-scale flow behavior in the form of particle agglomeration called clusters and streams. The size and frequency of these clusters


19

and streams are also highly dependent on operation and design conditions. Both the cluster size and frequency have a substantial impact on both catalytic reaction rates and heat transfer. Fortunately, a clearer understanding of these macro-scale and meso-scale behaviors are under intensive investigations. Today, fast-fluidized beds and risers can be designed with the solids concentration and radial profile in mind. However, this can only be achieved if one has an understanding of gas–solid hydrodynamics.

Macro-Scale Behavior: Solids Profile

Axial Profile of Solids in a Fast-Fluidized Bed and Riser In 1971, Reh noted that there exists an axial gradient of solids concentration in a riser similar to that observed in a fast-fluidized bed. Figure 14 illustrates these subtitle differences in the axial solids concentration profile commonly observed throughout a fast-fluidized bed and riser. As the gas velocity increases through a fluidized bed, the boundary between the dense fluidized-bed region and the lean freeboard region becomes indistinguishable. Indeed, having a distinguishable bed height is one of the indicators for fast fluidization (see the subsection Fast-Fluidized Regime). However, it was surprising that this axial gradient also exists in a riser (i.e., pneumatic conveying region) where higher gas velocities are used. This behavior was later quantified by Li and Kwauk [26], Weinstein et al. [40], Hartge et al. [41], and Rhodes et al. [42], who found that the axial gradient of the solids concentration exhibited an S-shaped curve. Horio [43] suggested that this S-shaped curve is restricted to units with a large L/D, riser length to diameter ratio. Large-scale units, such as atmospheric fluidized-bed combustors, may not exhibit this axial profile. Unfortunately, data on large units are limited and the solids concentration profiles in these units are still a subject of debate.

FIG. 14

Axial solids concentration profile in (a) a fast-fluidized bed and (b) a riser.

20


FIG. 15 Length-normalized pressure drop in a riser with increasing solids flux.

Recent studies have shown that the design of the entrance and exit region in a riser can have a substantial effect on the resulting performance. As illustrated in Fig. 15, the pressure drop, normalized by the distance between pressure taps, versus the riser length to diameter ratio, L/D, can be affected by the entrance design for 20 to 30 L/D’s. Furthermore, this effect becomes more pronounced with higher solids fluxes. Using an x-ray absorption technique, Kostazos et al. [44] were able to further substantiate this effect by examining the radial profile in a riser at varied feed ports. Their results showed that an asymmetric position of the feed manifested itself in the asymmetry of the radial profile at an axial position of up to 30 L/D’s.

Radial Profile of Solids in a Fast-Fluidized Bed and Riser The radial profile of solids that exists in fast-fluidized beds and risers is even more surprising. At some point beyond the entrance region of the fast-fluidized bed or riser, particles segregate toward the wall to form a core–annulus profile, as illustrated in Fig. 16. Studies using kinetic sampling probes, a γ-ray densitometer, and fiber optic probes were able to resolve this core–annulus profile [1,45–48]. Their results showed that the core consists of a lean concentration of solids moving up the riser, whereas the annulus consists of a dense concentration of particles. At moderate solids fluxes, particles in the annulus region actually exhibit a downward velocity, as shown in Fig. 16 [46,48,49]. Karri and Knowlton were able to quantify this downflow as a function of radial profile by measuring the solid mass fluxes in a 20-cm-diameter by 14-m-high riser [49]. Figure 17 presents their results where downflow in the annulus regions was observed for solid mass fluxes of 49 and 93


21

FIG. 16 Representation of the core–annulus profile in a riser where downflow is observed near the walls. (Adapted from Ref. 48.)

kg/m2 s. Miller and Gidaspow [50] showed that the largest magnitude of annular downflow flux at and near the wall was near the bottom of the riser. At less than 2 m from the inlet of Miller’s 7.5-cm-diameter riser, downward fluxes were several times the average feed flux [51]. The implications of this behavior can be substantial. For many catalytic reactions, backmixing near the feed region and, to a lesser extent, up throughout the riser can have a significant impact on productivity. Fortunately, many of these reactions require very high solids fluxes where downflow may be less of an issue. For example, Fig. 17 shows that operating Karri and Knowlton’s riser [49] at or above solid fluxes of 195 kg/m2 s, results in a core–annulus profile where particles at the wall move in the same direction as those in the core region (positive solids mass flux of ⬃1 kg/m2 s). In this case, backmixing was limited. Similar findings

FIG. 17 The effects of solids mass flux on the radial net solids mass flux profile. (Adapted from Ref. 49.)

22


FIG. 18 Representation of the core–annulus profile in a riser where upflow is observed near the walls. (Adapted from Ref. 49.)

have also been reported by Issangya et al. [52]. A representation of this behavior is presented in Fig. 18. Particle segregation appears also to be influenced by the core–annulus profile in a riser. Karri and Knowlton [49] observed that in the presence of downflow, the particle size distribution in the annulus region was larger than that found in the core. In contrast, this effect appears to only occur for downflow operations. Particle segregation was not observed for core–annulus upflow profiles for either very high or low solid mass fluxes [49]. Jones et al. [53] examined this phenomena using, the Laser Doppler Velocimetry (LDV) of particle-laden jets. Their results showed that eddies or recirculation zones were responsible for this particle segregation. Hence, the high shear and resulting recirculation zones generated from the solids downflow near the wall may be responsible for the segregation effect observed by Karri and Knowlton [49]. With upflow at the wall, the low shear may not generate strong enough eddies to effect the particle size distribution across the riser diameter. There are also design features that can reduce backmixing in risers. Baffles can induce wakes and turbulence, which limit the core–annulus profile. Of course, the added attrition caused by baffles needs to be factored into the design process. Another option is to use secondary feeds to produce a higher plug flow or uniform solids velocity profile at the entrance region. A core–annulus profile may still develop further up the riser, but backmixing is less severe in this region. As with the axial profile, the design of the entrance and exit region can have a substantial effect on the solid radial concentration profile. Rhodes et al. [42] used a nonisokinetic sampling probe to examine the radial solids loading in a 0.09-minner diameter by 7.2-m-high riser. Their results showed that a side solids feed resulted in a nonuniform radial distribution of solids beyond 40 L/D’s, as depicted in Fig. 19. In addition, Rhodes et al. noted that the asymmetries in solids radial distribution were more noticeable in the interphase between the dense and dilute regions. Thus, depending on the design of the feed region, a nonuniform radial profile may exist throughout many industrial risers. In a similar fashion, the exit configuration of a riser can have an impact on the solids profile for several L/D’s below the exit region. Brereton and Grace [54] observed this effect for smooth and abrupt riser exits. As shown in Fig. 20, using


23

FIG. 19 Illustration of the nonuniform solids radial profile in a riser due to solids feed on the side of the riser (not drawn to scale). (Adapted from Ref. 28.)

FIG. 20 Effect of exit configuration on solids volume fraction for a 0.15-m-diameter by 9.3-m-high riser with a superficial gas velocity of 7.1 m/s, initial solids flux of 73 kg/m2 s, and 148µm sand particles. (From Ref. 54).

24


FIG. 21 Solids flux ratio with respect to radial position for a riser with a smooth, rounded exit at a superficial gas velocity of 4.2 m/s, solids flux of 50 kg/m2 s, and 80-µm sand particles. (From Ref. 55.)

a smooth, wide-radius bend to terminate the riser resulted in little deviation in the axial solids concentration profile. However, an abrupt bend, such as a square bend or tee, resulted in backmixing, which affected the overall riser solids volume fraction profile up to 20 L/D’s below the exit region. Similar effects for solids fluxes are reflected in the data of Kruse and Werther [55] who compared normalized solid fluxes to radial solids loadings for a 0.4-mdiameter by 15.1-m-high riser, as shown in Figs. 21 and 22. For smooth bends, substantially less downflow is observed compared to the abrupt exit configurations. In addition, the region of downflow for the abrupt exit configuration was over twice the size of that observed for the smooth configuration. These results provide a good example of the importance of riser design for chemical production. For combustors, where backmixing is tolerable and sometimes even desired, asymmetric feed designs and abrupt exits are less critical. How-

FIG. 22

Solids flux ratio with respect to radial position for a riser with an abrupt, squared exit at a superficial gas velocity of 4.2 m/s, solids flux of 71 kg/m2 s, and 80-µm sand particles. (From Ref. 55.)


FIG. 23

25

Illustration of riser entrance region for a more uniform solids loading profile.

ever, in chemical production such as in oxidation and chlorination, asymmetric solids profiles and backmixing can seriously reduce selectivity and activity. Fortunately, both the entrance and exit can be designed such that minimal asymmetric solids profiles and backmixing ensues. For the entrance region, care needs to be taken such that entering solids are well mixed with the entraining gas. One such design is shown in Fig. 23. A fluidizing gas is used to distribute incoming solids, and one or more jets are used to entrain catalyst into the riser. Similarly, the exit region should have either a long radius bend or a disengagement section. Typically, industrial risers have a stripper section at the top of the riser to not only strip gas but also minimize exit effects on the riser, as shown in Fig. 24.

Meso-Scale Behavior: Clusters and Streamers Riser sections in circulating fluidized beds exhibit a core–annulus profile with downward flow resulting in the formation of clusters and streamers of particles.

FIG. 24 Illustration of riser exit region (i.e., stripper) for a more uniform solids loading profile.

26

Circulating Fluidized Bed Reactors This phenomenon was proposed by Squires et al. [56] and Yerushalmi and Avidan [21,57]. Under the assumption that the pressure drop equals the weight of the solids in suspension, the resulting slip velocity, calculated as v slip ⫽ v g ⫺

Gs ρs εs

(34)

was found to be several times larger than the terminal velocity [51]. Because the slip velocity cannot exceed the terminal velocity, it was postulated that the particles must be forming clusters that effectively act as larger particles. Today, cluster and streams are frequently observed. High-speed movies [24], laser sheet [58], infrared imagining [59], and fiber-optic probes [60–62] all reveal the presence of wave packet of particles near the riser wall moving with the downward annulus flow in continuous but dynamic and unstable sheets. These sheets of particles (commonly called clusters, streamers, swarms, strings, or strands) are represented in Fig. 25. Laser sheet and fiber-optic studies of Horio [43] and Rudnick and Werther [61] have further demonstrated that these clusters are three dimensional in nature and can be found in the annulus and core regions. In general, it was observed that clusters move in a direction parallel to the flow of the suspension phase. In other words, in a core–annulus profile with downward flow at the walls, clusters in the annulus region flow downward and clusters in the core region flow upward. Soong et al. [63] experimentally measured the cluster length and time-averaged local solid volume fraction in riser flow. Their results were in agreement with Yerushalmi and Avidan’s [57] earlier empirical correlation of d c ⫹ d p ⫹ ε s (0.027 ⫺ 10 d p ) ⫹ 32 ε 6s

(35)

However, Soong et al. replaced the average local solids volume fraction of particles, ε s, by the solids volume fraction of a cluster, ε cl, as d c ⫹ d p ⫹ ε cl (0.027 ⫺ 10d p ) ⫹ 32 ε 6cl

FIG. 25 An illustration of the cluster or streamers often observed in risers.

(36)


27

Gu and Chen [64] further correlated Soong’s data such that the solids volume fraction of a cluster can be related to the local solids volume fraction of particles using the expression

冤冢ε 冣冥

ε cl ⫽ ε s,max 1 ⫺

εs

3.4

(37)

s,max

where the maximum local solids volume fraction, ε s,max, is 0.57. Tsuo and Gidaspow [65] observed that, for Group A powders, the clusters were 2–3 cm in length in the down-flowing annulus region. Furthermore, cluster density increased with increasing solids flux, decreasing gas velocity, or decreasing pipe diameter. The mechanism for the formation and degeneration of clusters in a riser is still under some dispute. Tsuo and Gidaspow [65] proposed that clusters are the result of partially inelastic collision with the walls. Particles hit the wall, lose energy, and fall, to collide with another particle. This process continues until a cluster is formed. Senior and Grace [66] proposed that inelastic wall collisions cannot account for all the energy loss needed to form clusters. For this to happen, wall collisions would need a coefficient of restitution of less than 0.1, which is unlikely. Perfectly elastic collisions have a coefficient of restitution of 1. Instead, Senior and Grace proposed that the balance between shear-induced lift and drag forces on a particle act to momentarily detain particles at the wall region of the riser. After a particle collides with the wall, it has insufficient momentum to counteract the lift force. As a result, the particle continues to hit the wall, each time losing more lateral velocity. When the particle-to-wall friction slows the particle below the local gas velocity, lift forces acting in the opposite direction move the particle away from the wall to the near-wall regions. Clusters form when many particles undergo this lateral-velocity-reduction process. The downward motion of a cluster is caused by the net lift forces on a cluster being less than the sum of the forces on each individual particle. Particle migration to the wall is not only dependent on axial and lateral velocities but also on the particle diameter. Lee and Durst [67] found that 100- and 200µm-diameter glass beads readily accumulated at the wall, whereas larger particles, with a 400–800-µm diameter, did not. The larger particles were traveling at significantly lower axial velocities and were less influenced by lift forces directed to the wall. Tsuji et al. [68] was also able to measure this crossover of particles to the wall for smaller particles. Senior and Grace [66] were able to model these trajectories and found similar conclusions. For particles larger than 500 µm in diameter, no range of lateral velocities was found that slowed the particle significantly enough for crossover to the wall. Yet, for 230-µm particles, a significant concentration of particles was predicted to accumulate near the wall for initial lateral velocities of 4.5–5.5 cm/s. For 40-µm particles, concentrations were found to be an order of magnitude larger at the wall than that found for 230-µm particles with lateral velocities ranging from 3 to 22 cm/s. What is interesting here is that if particle collisions with the wall help create the formation of clusters, how do cluster form in the core region as observed by Horio [43] and Rudnick and Werther [61]. Furthermore, Karri and Knowlton [49]

28

Circulating Fluidized Bed Reactors observed that particles in the annulus region with a downward flow had a larger particle size distribution than that in the core. In a core–annulus profile where both regions have an upflow profile, no particle size distribution effects were observed. Both of these results contradict the above postulate of Senior and Grace in which a wall is needed for cluster formation and larger particles prefer the core region. Most likely, the magnitude of the solids flux or solids concentration may have a significant impact on the locality of cluster formation and particle-size-segregation effects. These macroscopic properties may also need to be considered. Horio [43] proposed that another mechanism may be responsible for cluster formation. A particle in flight can have either attractive or repulsive forces with nearby particles. Two particles traveling perpendicular to the gas flow tend to repeal each other, whereas two particles aligned parallel with the flow tend to attract each other due to the nearest-neighbor effect on lift and drag. Yet, in riser flow, particle alignment is not stable, as particles undergo collisions, bumping, tumbling, and other nonelastic processes. This may be the very mechanism to ensure clustering. The combination of nonelastic processes and the parallel-aligned particle flow may provide the attractive force needed to promote clustering. Furthermore, Horio noted that buoyancy forces dominate in lean-phase regions, whereas gravitational forces dominate in dense-phase regions. This results in a shear between clusters and the lean-phase region. The interaction of these forces controls the development of particle groups to form steady but turbulent structures of a certain characteristic length. Two-dimensional simulations of particles show this very event, where a homogeneous suspension evolved into clusters [69].

Closing Remarks To date, a complete description of the CFB hydrodynamics is not possible even after decades of intensive research. Hence, predicting accurate pressure profiles, reaction productivities, heat transfer, and gas and solid residence times is still more art than science. Yet, circulating-fluidized-bed reactors have been in productive use for more than 50 years. Indeed, a broad knowledge base exists for the design of fluidized catalyst crackers using empirical-based models. Even though a fundamental understanding of the physics behind gas–solid flow is limited, the design, construction and operation of these units can be done in a relatively short period of time. For the design, construction, and operation of CFBs in the chemical industry, this is not the case. Unlike the petroleum industry where FCC catalysts have similar properties, the physical properties of the catalyst commonly used in the chemical industry are diverse. Thus, it is more important to understand and apply at least the known physics in gas–solid flow. This is not a daunting task; just careful consideration of each design aspect of the CFBs needs to be addressed.


29

Nomenclature Ar A surface db d b,max dc dp d p′ d p,ave d *p D D L-valve g gc Gs L Lb L L-valve Lu P 45µm Q aeration Qg r R Re p Re pmf ug u mc u mb u mf u mr u ms u mt ui u0 u* ut V˙ g v slip z ε ε bp ε cl

Archimedes’ number Surface area Bubble diameter Equilibrium bubble diameter Cluster diameter Particle diameter Effective particle diameter Average particle diameter Dimensionless particle diameter Vessle diameter Diameter of an L-valve Acceleration of gravity Newton’s law proportionality factor, 9.8 m/s2 or 32.2 ft/s2 Solids flux rate Length Height of fluidized bed Height of L-valve (distance between L-valve feed and horizontal leg into reactor) Height of fluidized-bed vessel Proportion of particles with a diameter less than 45 µm Volumetric flow rate of the gas for L- or J-valve Volumetric flow rate of the gas Radius Vessel radius Particle Reynolds number Particle Reynolds number at minimum fluidization Superficial gas velocity Superficial gas velocity at the onset of pneumatic conveying in a fluidized bed Superficial gas velocity at the onset of bubbling in a fluidized bed Superficial gas velocity at minimum fluidization Superficial gas velocity at the onset of fast fluidization in a fluidized bed Superficial gas velocity at the onset of slugging in a fluidized bed Superficial gas velocity at the onset of the turbulence regime in a fluidized bed Inlet velocity of a cyclone Tangential velocity at the wall of a cyclone Dimensionless gas velocity Particle terminal velocity Volumetric gas flow rate Slip velocity Distance Solids void fraction Solids void fraction for a packed bed Average solids void fraction of a cluster

30

Circulating Fluidized Bed Reactors ε ms ε mf εs ∆P λp µ ρg ρs Φ

Solids void fraction at the onset of pneumatic conveying in a fluidized bed Solids void fraction for a bed at minimum fluidization Average solids void fraction Pressure drop Dimensionless solids friction coefficient Gas viscosity Gas density Solids density Sphericity

Appendix A: Geldart Powder Classification Simpson and Rodger [70], Jackson [71], and Verloop and Heertjes [72] suggested that the fluidization of particles can be classified into two categories: particulate and aggregative. The particulate category pertains to powders that fluidize in a liquidlike fashion. As the superficial gas velocity increases, these particles move further apart in an independent fashion. In contrast, particles in the aggregates catagory would exhibit bubble formation with increasing gas velocities. For the aggregate particles, the resulting dense phase remains unchanged in terms of solids concentrations after initial fluidization. Most particles adhere to the aggregative behavior. Unfortunately, this two-category methodology falls short in adequately describing fluidization behavior. The particulate category is limited to a small group of particles leaving the aggregative category to describe everything else. To address this shortcoming, Geldart and Rhodes [1,73] demonstrated that particles could be classified into four distinct categories or groups. Today, these groups are referred to as Geldart Group A through D. Figure A1 illustrates how the fluidization behavior of a particular powder can be predicted using the Geldart Group classification. This graph is the basic foundation of modern-day fluidization engineering. By comparing the particle density (less the gas density) with the mean particle diameter, the ‘‘type’’ of fluidization can be determined. In general, Group A powders undergoing fluidization behave significantly different than the other groups. Thus, the design and operation of a fluidized-bed unit containing a Group A powder would not be the same as that used for a Group B powder. To better clarify these differences, a description of the fluidization behavior is presented in the order of increasing particle size. Geldart Group C powders are typically less than 50 µm in diameter and are the most difficult to fluidize. These particles are considered cohesive and almost always exhibit significant channeling during fluidization. To limit this effect, Group C powders are usually fluidized with the aid of baffles and/or mechanical vibration. Sometimes, larger particles, such as Group B powders, are added to the bed to promote smoother fluidization. Geldart Group A powders are the most common type of powder used in fluidization. For example, most FCC units are designed for Group A powders, which


31

FIG. A1 Geldart powder classification at ambient conditions. (Adapted from Ref. 7.)

typically have a mean particle diameter of 75 µm and a particle density of 1.0 g/ cm3. At low gas velocities, Group A powders tend to exhibit significant bed expansion without the formation of bubbles (i.e., smooth fluidization). At higher gas velocities (i.e., greater than u mb ), bubbles appear and rise more rapidly than the gas in the rest of the bed. Gas in the dense or emulsion phase tend to percolate through the bed compared to the residence time of the gas in the bubbles. Typically, Group A powders do not promote maximum bubble size greater than 10 cm [7]. Geldart Group B powders have particle diameters typically ranging from 200 to 800 µm. Unlike Group A powders, where smooth fluidization is observed, these powders exhibit the formation of bubbles at the onset of fluidization. The bubble size in Group B powders can be large, on the order of feet in some cases. Group B powders fluidize easily and are used in a wide range of fluidization unit operations with little difficulties. Care should be taken that slugging does not occur in smaller fluidized beds. Geldart Group D powders have the largest particle diameters of all other Geldart groups. As a result, gas requirements for fluidization are large. These powders are typically processed in spouting beds where gas requirements are less than that needed in standard fluidized beds. During fluidization, Group D powders have enormous bubble diameters and slugging is commonly observed even in large fluidized beds. It should be noted that Figure A1 was developed for particles at ambient conditions. Under high pressures, a Group B powder may behave as a Group A powder Furthermore, the transition from one group to another is not well defined. In some cases, the behavior of a powder can be classified under more than one group. For example, some powders fluidize well as a Group A powder, but become permanently defluidized as a Group C powder in nonaerated horizontal sections of a CFB. These powders are sometimes referred to as Group AC powders.

32


Appendix B: Terminal Velocity The velocity of a free-falling particle will increase according to the acceleration of gravity. At some point, the drag on the particle will reduce the acceleration to zero and the particle reaches its terminal velocity. Haider and Levenspiel [74] defined the terminal velocity of a single particle as

冢

⫺1

冣

2.335 ⫺ 1.744Φ 18 ⫹ u *t ⫽ 2 (d*p ) (d*p )0.5

(B1)

where u*t ⫽ u t

冢

ρ 2g µ(ρ s ⫺ ρ g )g

冣

1/3

Thus, to avoid significant carryover in a fluidized bed, the superficial gas velocity should be less than the terminal velocity, u t.

References

1. D. Geldart, ‘‘Introduction,’’ in Gas Fluidization Technology (D. Geldart, ed.), John Wiley & Sons, Chichester, 1986, p. 5. 2. R. H. Perry and C. H. Chilton, Chemical Engineers’ Handbook, 5th ed., McGrawHill, New York, 1973, p. 20–71. 3. M. Sittig, Chem. Eng., 60, 219–231 (1953). 4. K. Weissermel and H.-J. Arpe, Industrial Organic Chemistry, 2nd ed., VCH, Weinheim, 1993, p. 302. 5. J. L. Callahan, E. C. Milberger, and R. K. Grasselli, ‘‘Process for Preparing Olefinically Unsaturated Aldehydes and Nitriles,’’ U.S. Patent No. 3,4827,343 (February 11, 1969). 6. J. L. Callahan, H. F. Hardman, and E. C. Milberger, ‘‘Reactor for Contacting Gases and a Particulate Solid,’’ U.S. Patent No. 4,152,393 (May 1, 1979). 7. D. Kunni and O. Levenspiel, Fluidization Engineering, 2nd ed., Butterworth–Heinemann, Boston, 1991, p. 44. 8. W. L. McCabe and J. C. Smith, Unit Operations of Chemical Engineering, 3rd ed., McGraw-Hill, New York, 1976, p. 149. 9. C. Y. Wen and Y. H. Yu, AIChE J., 12, 610–612 (1966). 10. J. F. Richarson, ‘‘Incipient Fluidization and Particle Systems,’’ in Fluidization (J. F. Davidson and H. Harrison, eds.), Academic Press, New York, 1971, p. 26–64. 11. S. C. Saxena and G. J. Vogel, Trans. Inst. Chem. Eng., 55, 184–189 (1977). 12. S. P. Babu, B. Shah, and A. Talwalkar, AIChE Symp. Ser. 176, 176–186 (1978). 13. J. R. Grace, ‘‘Fluidized Bed Hydrodynamics,’’ In Handbook of Multiphase Systems (G. Hetsroni, ed.), Hemisphere, Washington, DC, 1982, p. 8–1. 14. D. C. Chitester, R. M. Kornosky, L.-S. Fan, and J. P. Danko, Chem. Eng. Sci., 39, 253–261 (1984).


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15. A. R. Abrahamsen and D. Geldart, Powder Technol., 26, 47–55 (1980). 16. J. F. Davidson, and D. Harrison, Fluidized Particles, Cambridge University Press, Cambridge, 1963. 17. P. S. B. Stewart and J. F. Davidson, Powder Technol., 1, 61–80 (1967). 18. J. Yerushalmi and N. T. Cankurt, Powder Technol., 24, 187–205 (1979). 19. J. R. Grace and H. Bi, ‘‘Introduction to Circulating Fluidized Beds,’’ in Circulating Fluidized Beds (J. R. Grace, A. A. Avidan, and T. M. Knowlton, eds.), London: Blackie Academic & Professional, London, 1997, pp. 1–20. 20. H. T. Bi and J. R. Grace, Chem. Eng. J., 57, 261–271 (1995). 21. J. Yerushalmi, ‘‘High Velocity Fluidized Beds,’’ in Gas Fluidization Technology (D. Geldart, ed.), John Wiley & Sons, New York, 1986, pp. 127–196. 22. M. G. Schnitzlein and H. Weinstein, Chem. Eng. Sci., 43, 2605–2614 (1988). 23. W. C. Yang, Powder Technol., 35, 143–150 (1983). 24. L. Reh, Chem. Eng. Prog., 67, 58–63 (1971). 25. J. Werther, Chem. Eng. Sci., 35, 372–379 (1980). 26. Y. Li and M. Kwauk, ‘‘The Dynamics of Fast Fluidization,’’ in Fluidization (J. R. Grace and J. M. Matsen, eds.), Plenum, New York, 1980, pp. 537–544. 27. A. A. Avidan and J. Yerushalmi, Powder Technol., 32, 223–232 (1982). 28. M. J. Rhodes, Chem. Eng. Res. Des., 67, 30–37 (1989). 29. L. S. Leung, Powder Technol., 25, 185–190 (1980). 30. G. E. Klinzing, Gas–Solid Transport, McGraw-Hill, New York, 1981. 31. J. R. Grace, Can. J. Chem. Eng., 64, 353–363 (1986). 32. F. A. Zenz and D. F. Othmer, Fluidization and Fluid Particle Systems, Van Nostrand Reinhold, New York, 1960. 33. J. J. van Deemter, in Fluidization, (J. R. Grace and J. M. Matsen, eds.), Plenum, New York, 1980, pp. 69–89. 34. M. Horio, N. M. Hoshiba, K. Morishita, Y. Kobukai, J. Naito, O. Tachibana, K. Watanabe, and N. Yoshida, ‘‘Coal Combustion in a Transparent Circulating Fluidized Bed,’’ in Circulating Fluidized Bed Technology (P. Basued, ed.), Pergamon, New York, 1986, pp. 225–262. 35. N. M. Catipovic, G. N. Jovanovic, and T. J. Fitzgerald, AIChE J., 24, 543–547 (1978). 36. E. Muschelknautz and V. Grief, ‘‘Cyclones and Other Gas–Solids Separators,’’ in Circulating Fluidized Beds (J. R. Grace, A. A. Avidan, and T. M. Knowlton, eds.) Blackie Academic & Professional, London, 1997, pp. 181–213. 37. W. C. Yang and T. M. Knowlton, Powder Technol., 77, 49–54 (1993). 38. D. R. Jones and J. F. Davidson, Rheol. Acta, 4, 180–192 (1965). 39. J. Sinclair and T. Mallo., FED (Am. Soc. Mech. Eng.), 245, 12/58–12/63 (1998). 40. H. Weinstein, R. F. Graff, M. Meller, and M. J. Shao, ‘‘The Influence of the Imposed Pressure Drop Across a Fast Fluidized Bed,’’ in Fluidization IV (K. Kunii and R. Toei, eds.), Engineering Foundation, New York, 1983, pp. 299–306. 41. E. U. Hartge, D. Rensner, and J. Werther, ‘‘Solid Concentration and Velocity Patterns in Circulating Fluidized Beds,’’ in Circulation Fluidized Bed Technology II (P. Basu and J. F. Large, eds.), Pergamon Press, Oxford, 1988, pp. 165–180. 42. M. J. Rhodes, M. Sollaart, and X. S. Wang, ‘‘Structure of the Dense–Dilute Interface in Fast Fluidization,’’ in Fluidization IX (L-S. Fan and T. M. Knowlton, eds.), Engineering Foundation, New York, 1998, pp. 141–148. 43. M. Horio, ‘‘Hydrodynamics,’’ in Circulating Fluidized Beds (J. R. Grace, A. A. Avidan, and T. M. Knowlton, eds.), Blackie Academic & Professional, London, 1997, pp. 21–85. 44. A. E. Kostazos, H. Weinstein, and R. A. Graff, ‘‘The Effect of the Location of Gas Injection on the Distribution of Gas and Catalyst in a Riser,’’ in Fluidization IX (L-S. Fan and T. M. Knowlton, eds.), Engineering Foundation, New York, 1998, pp. 221–228.

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Circulating Fluidized Bed Reactors 45. H. Weinstein, M. Shao, M. Schnitzlein, and R. A. Graff, ‘‘Radial Variation in Void Fraction in a Fast Fluidized Bed,’’ in Fluidization V (K. Ostergarrds and A. Sorensen, eds.), AIChE Proceedings on the 5th Engineering Foundation Conference on Fluidization, New York, 1986, p. 329. 46. R. Bader, J. Findlay, and T. M. Knowlton, ‘‘Gas–Solid Flow Pattern in a 30.5 cm Diameter Circulating Fluidized Bed,’’ in Circulating Fluidized Bed Technology II (P. Basu and J. F. Large, eds.), Pergamon Press, New York, 1988, pp. 123– 127. 47. P. A. Galtier, R. J. Pointer, and T. E. Patureaux, ‘‘Near Full-Scale Cold Flow Model for the R2R Catalytic Cracking Process,’’ in Fluidization VI (J. R. Grace, L. W. Shemilt, and M. A. Bergougnou, eds.), Engineering Foundation, New York, 1989, pp. 17–24. 48. A. Miller, ‘‘Dense, Vertical Gas-Solid Flow in a Pipe,’’ Ph.D. dissertation, Illinois Institute of Technology, Chicago (1991). 49. R. S. B. Karri and T. M. Knowlton, ‘‘Flow Direction and Size Segregation of Annulus Solids in a Riser,’’ in Fluidization IX (L-S. Fan and T. M. Knowlton, eds.), Engineering Foundation, New York, 1998, pp. 189–196. 50. A. Miller and D. Gidaspow, AIChE J., 38, 1801–1815, (1992). 51. D. Gidaspow, Multiphase Flow and Fluidization: Continuum and Kinetic Theory Descriptions, Academic Press, Boston, 1994, p. 209. 52. A. S. Issangya, D. Bai, and J. R. Grace, ‘‘Solids Flux Profiles in a High Density Circulation Fluidized Bed Riser,’’ in Fluidization IX (L-S. Fan and T. M. Knowlton, eds.), Engineering Foundation, New York, 1998, pp. 196–204. 53. N. E. Jones, C. U. Yurteri, and J. L. Sinclair, ‘‘Effects of Solids Loading on Particle Motion in Gas–Solid Flow,’’ AIChE Annual Meeting Preprint ‘‘Fluidization and Fluid–Particle Systems,’’ in Dallas, TX (L. R. Glicksman, ed.), 1999, p. 65. 54. C. Brereton and J. R. Grace, ‘‘End Effects in Circulation Fluidized Bed Hydrodynamics,’’ in Circulating Fluidized Bed Technology IV (A. A. Avidan, ed.), AIChE, New York, 1990, pp. 137–144. 55. M. Kruse and J. Werther, J. Chem. Eng. Process., 34, 185–203 (1995). 56. A. M. Squires, M. Kwauk, and A. A. Avidan, Science, 230, 1329–1337 (1985). 57. J. Yerushalmi and A. A. Avidan, Powder Tech., 32, 223–232 (1982). 58. M. Horio and H. Kuroki, Chem. Eng. Sci., 49, 2413–2421 (1994). 59. P. D. Noymer and L. R. Glicksman, ‘‘Temperatures of Clusters at the Wall of a Circulating Fluidized Bed,’’ in Fluidization IX (L-S. Fan and T. M. Knowlton, eds.), Engineering Foundation, New York, 1998, pp. 429–436. 60. H. T. Ishii, T. Nakajima, and M. Horio, J. Chem. Eng. Japan, 22, 484–490 (1989). 61. C. Rudnick and J. Werther, ‘‘The Discrimination of Cluster Characteristics from Fiber-Optical Probe Signals in Circulating Fluidized Beds,’’ in Fluidization IX (L-S. Fan and T. M. Knowlton, eds.), Engineering Foundation, New York, 1998, pp. 572– 580. 62. S. Krol and H. de Lasa, ‘‘CREC–GS–Optiprobe for Particle Cluster Characterization,’’ in Fluidization IX, (L-S. Fan and T. M. Knowlton, eds.), Engineering Foundation, New York, 1998, pp. 565–572. 63. C. H. Soong, K. Tuzla and J. C. Chen, ‘‘Experimental Determination of Cluster Size and Velocity in Circulation Fluidized Bed,’’ in Fluidization VIII, (J. F. Large and L. Claude, eds.) Engineering Foundation, New York, 1995, pp. 219–227. 64. W. K. Gu and J. C. Chen, ‘‘A Model for Solid Concentration in Circulating Fluidized Beds,’’ in Fluidization IX (L-S. Fan and T. M. Knowlton, eds.), Engineering Foundation, New York, 1998, pp. 501–508. 65. Y. P. Tsuo and D. Gidaspow, AIChE J. 36, 885–896 (1990). 66. R. C. Senior and J. R. Grace, ‘‘Formation of Particle Streamers in the Wall of Circulat-

Pressure and Temperature in Fluidized Beds

67. 68. 69.

70. 71. 72. 73.

74.

35

ing Fluidized Bed Risers,’’ in Fluidization IX (L-S. Fan and T. M. Knowlton, eds.), Engineering Foundation, New York, 1998, pp. 165–172. S. L. Lee and F. Durst, Int. J. Multiphase Flow, 8, 125–146 (1982). Y. Tsuji, Y. Morikawa, and H. J. Shiomi, J. Fluid Mech., 139, 417–434 (1984). S. Yonemura, T. Tanaka, and Y. Tsuji, ‘‘Cluster Formation in Dispersed Gas–Solid Glow: Effects of Physical Properties of Particles,’’ Proceedings to 2nd International Conference on Multiphase Flow, Kyoto, 1995, vol. 3, pp. PT4-25–PT4-30. H. C. Simpson and B. W. Rodger, Chem. Eng. Sci., 16, 153–180 (1961). R. Jackson, Trans. Inst. Chem. Eng., 41, 13–28 (1963). J. Verloop and P. M. Heertjes, Chem. Eng. Sci., 25, 825 (1970). D. Geldart and M. J. Rhodes, ‘‘From Minimum Fluidization to Pneumatic Transport: A Critical Review of the Hydrodynamics,’’ in Circulation Fluidized Bed Technology (P. Basu, ed.), Pergamon Press, New York, 1986, pp. 21–31. A. Haider and O. Levenspiel, Powder Technol. 58, 63–70 (1989). RAY A. COCCO

Effect of Pressure and Temperature in Bubbling Fluidized Beds

Introduction Typical commercial gas fluidized-bed processes operate at up to 7.1 ⫻ 106 Pa (⬃70 atm) pressure and 1273 K (1000°C). Hence, an understanding of the effect of pressure and temperature on hydrodynamics is critical for propert fluidized-bed system design. Temperature and pressure affect fluidized-bed systems because they affect the gas density and gas viscosity, two critical parameters impacting bed operation. These effects cannot be considered independently of particle size, as particle size dramatically impacts fluid–particle interactions. In this section, we will summarize the impact of pressure and temperature on the critical hydrodynamic parameters of minimum fluidization velocity, u mf, bed voidage at minimum fluidization velocity, ε mf, minimum bubbling velocity, u mb, bubble diameter, d b, bubble frequency, f, dense phase voidage, ε D, and particle entrainment and elutriation rates, K *i ∞. We will also summarize the reported effect of pressure on chemical reaction conversion because reactor performance is the result of all combined effects the pressure and temperature may have on hydrodynamics as well as on the reaction. The reader is referred to two in-depth reviews of the subject [1,2].

36


Two-Phase Flow Model A bubbling fluidized bed can be considered as consisting of two phases (Fig. 1): (1) a dense (or emulsion) phase of solid particles with a surrounding interstitial dense (or emulsion) phase gas and (2) a dilute phase consisting of gas jets which protrude up through a perforated gas distributor or series of nozzles, and gas bubbles that form at the top of the jets. The gas that enters a fluidized bed at the volumetric flow rate q G(in) enters in the form of gas jets. At the top of the jets, this gas is distributed between bubbles that form (q B) and the interstitial dense phase gas (q D) that flows through the solid particles. Mass and heat transfer (interchange) occur between the jets and the dense phase gas, between the bubbles and the dense phase gas, and between the dense phase gas and solids. Heat is also transferred by the dense phase to internal heat-exchange surfaces. A critical factor affecting the overall performance of a fluidized bed is the interchange between the dilute and dense phases. Interchange rates can be increased by (1) increasing the ratio of the gas jet volume to the bubble volume in the bed (i.e., VJ /VB), (2) increasing the ratio of the gas flowing through the dense phase to the gas flowing as bubbles (i.e., q D /q B), and (3) maintaining a small bubble size throughout the bed. Small bubbles provide higher mass-transfer rates and increased residence time in the bed. It is imperative to note that although we want bubbles to be small, bubbles must not disappear. Bubbles must be present in order to achieve rapid mixing in the reactor, contributing to high heat-transfer rates to internal surfaces.

FIG. 1 Transport, phases, and pressure in fluidized beds.


37

Pressure and temperature can have a significant effect on the distribution of gas between the bubble and dense phases and the mass- and heat-transfer rates.

Minimum Fluidization Conditions The minimum fluidization velocity, u mf, is a fundamental characteristic of a fluidized bed. Its accurate prediction is important for the successful design and operation of a fluidized-bed process. At minimum fluidization conditions, the drag force by upward-moving gas equals the weight of particles lifted in the fluidized bed. The following relationship can be shown to hold (see the Nomenclature section for terms): Ar ⫽ 150

1.75 1 ⫺ ε mf Re mf ⫹ 3 Re 2mf 2 3 φ ε mf φε mf

(1)

2C 1 1 Re mf ⫹ Re 2mf C2 C2

(2)

or Ar ⫽ where Ar ⫽

d 3p ρ g(ρ p ⫺ ρ g)g , µ 2g C1 ⫽

d p u mf ρ g

Re mf ⫽

µg

42.86(1 ⫺ ε mf) φ

(3)

φε 3mf 1.75

(4)

C2 ⫽

Working on the assumption that the particles constituting the bed can be approximated by a constant value of φ and that ε mf remains constant over the entire range of operating conditions of temperature and pressure, various investigators have proposed values of C 1 and C 2 on the basis of experimental results to be in the range 18.75–33.7 and 0.0313–0.0651, respectively [3]. The first terms on the right-hand sides of Eqs. (1) and (2) are important if laminar, or viscous, flow predominates in the system, whereas the second terms are important if turbulent, or inertial, flow predominates. For small particles (Re mf ⬍ 20), the simplified form of Eqs. (1) and (2) is u mf ⫽

冢冣冢 C2 2C 1

冣

d 2p (ρ p ⫺ ρ g)g µg

(5)

38

Pressure and Temperature in Fluidized Beds For large particles (Re mf ⬎ 1000), the simplified form becomes u 2mf ⫽ C 2

冢

冣

d p (ρ p ⫺ ρ g)g ρg

(6)

With gas viscosity, µ g, almost independent of pressure and because ρ p ⬎⬎ ρ g for most materials, Eq. (5) indicates that u mf is relatively unaffected by changes in pressure for fine particles. On the other hand, according to Eq. (6), u mf for larger particles will vary with (1/ρ g)0.5, indicating that u mf decreases for increasing pressure. For changes in temperature, Eq. (5) indicates that u mf varies with 1/µ g . Because the gas viscosity, µ g, increases with temperature, u mf decreases with an increase in temperature for fine-particle systems in which viscous forces dominate. For large particles, Eq. (6) indicates that u mf will increase with temperature because an increased temperature results in a decreased ρ g. For systems in the intermediate regime (20 ⬍ Re mf ⬍ 1000), Eq. (2) can be rearranged and used: u mf ⫽

冢冣

µg [(C 21 ⫹ C 2 Ar)0.5 ⫺ C 1] dpρg

(7)

The trends that u mf is relatively unaffected by changes in pressure and decreases for increases in temperature for fine-particle systems and that u mf decreases for increases in pressure and increases for increases in temperature for large-particle systems are consistent with those reported experimentally. However, the absolute values of the predictions may be incorrect because of the difficulty in determining a representative value for d p and φ, or estimating ε mf. In the absence of reliable data, use C 1 ⫽ 33.7 and C 2 ⫽ 0.0408 [4] for fineparticle systems and C 1 ⫽ 28.7 and C 2 ⫽ 0.0494 [5] for coarse-particle systems. A recommended method for improving the accuracy of u mf is to first determine u mf experimentally at ambient conditions and to back-calculate an effective particle diameter (i.e., d eff) from Eq. (5), (6), or (7), using fixed C 1 and C 2. Using this effective particle diameter, calculate u mf at the desired conditions of temperature and pressure. This method substitutes an effective value for d p and φ, which are independent of temperature and pressure; however, it does not account for any changes in ε mf which might occur for changes in temperature and pressure. The effect of temperature and pressure on ε mf has been studied by a number of investigators. Several studies [5,6] have indicated that pressure has essentially no effect on ε mf for fine particles and a slight increasing effect on ε mf for larger particles. The effect of temperature on ε mf has been reported to be much more significant [7–10] and to affect fine-particle systems more than coarse-particle systems. The temperature effect appears to be the result of interparticle forces which affect packing properties [8,10]. The dependence of ε mf on temperature can be expressed in linear form as [8] ε mf ⫽ ε mf(amb) ⫹ k(T ⫺ T (amb) ) where the parameter k is a function of particle properties.

(8)


39

Although substantial data have been reported in the literature, no reliable correlations are available for predicting ε mf for a given system. However, it is clear that incorporating an accurate value of ε mf with an effective particle size, d eff, in Eq. (1) will allow an accurate calculation of u mf. Without experimental data, the use of Eq. (7) with the recommended values for C 1 and C 2 is suggested for calculating u mf.

Minimum Bubbling Velocity, umb and Dense Phase Voidage, εD As the gas velocity is increased above that required to incipiently fluidize the bed of particles, gas bubbles eventually form (at u 0 ⫽ u mb) and rise through the dense phase. In fine-particle systems at high pressures, it is possible to observe a particulately (or homogeneously) fluidized bed without bubbles for intermediate gas velocities between u mf and minimum bubbling, u mb (i.e., for u mf ⬍ u 0 ⬍ u mb). The minimum gas velocity at which bubbles appear, u mb, has been found to equal u mf for coarse-particle systems [11]. However, for fine-particle systems, there is a range of velocities between u mf and u mb over which the bed expands uniformly. The velocity range over which this ‘‘delayed bubbling’’ occurs can be extended with an increase in operating pressure. Abrahamsen and Geldart [12] observed that u mb and, hence, the region of particulate expansion is increased by adding fines (increased F, weight fraction of particles ⬍45 µm) to a fluidized bed of fine Group A powders: µ 0.523 exp(0.716F)u mf u mb 2300ρ 0.126 g g ⫽ 0.8 0.934 u mf d p g (ρ p ⫺ ρ g)0.934

(9)

Equation (9) indicates that u mb will increase for an increase in gas temperature and pressure (via increased u g and ρ g) and that the sensitivity to viscosity is more significant than the sensitivity to pressure. This equation has been shown to be valid over a wide range of pressures for fine-particle systems [13]. For fluidized beds of fine Group A powders that particulately fluidize (i.e., u mb ⬎ u mf), the voidage of the dense phase, ε D, exceeds ε mf. This voidage has been shown to be adequately described for high-pressure systems of fine powders [14] by the empirical correlation of Kmiec [15]: εD ⫽

(18Re ⫹ 2.7Re1.687)0.209 Ar 0.209

(10)

It is important to note that an increase in ε D results in a decrease in dense phase viscosity, µ D. The effect of u mb ⬎ u mf and ε D ⬎ ε mf is to increase the interchange between feed gases and solid particles because more feed gas flows directly through the dense phase (i.e., ⬎ q D /q B). Hence, overall reaction rates for catalytic and gas– solid fluidized-bed reaction processes involving fine particles will increase with

40

Pressure and Temperature in Fluidized Beds increased pressure due to the effect of pressure on fluidization hydrodynamics alone.

Jetting Region As gas flows through a perforated or multinozzle distributor plate, it enters a fluidized bed in the form of gas jets. This jetting region (Refs. 16 and 17, among others) has high gas–solid contacting efficiencies relative to the bubbling region above. Several investigators [18,19] have shown that the penetration of the gas jets into the fluidized bed increases substantially with an increase in operating pressure. The effect of gas viscosity was not found to be significant [20]. The contribution of pressure and temperature impacting the jet length, h j, is through their effect on gas density via the equation [19]:

冢冣冢冣

ρg hj ⫽ 21.2Fr 0.37 Re0.05 do ρp

0.68

dp do

0.24

(11)

Bubbling Region Bubble size is one of the most important parameters of conventional gas–solid fluidized beds. Overall reaction rates depend on the size of bubbles because bubble size is a primary contributor to bubble rise velocity, gas rate of exchange between phases, heat transfer, and fine-particle elutriation.

Fine-Particle Systems For fine-particle Group A systems, increased pressures yield smaller-sized bubbles. Studies have indicated [21] that the smaller bubbles result from a decrease in the stability of bubbles leading to their breakup into smaller voids. X-rays [22] of gas bubbles in fluidized beds have shown that an initial disturbance, indenting the upper surface of a bubble, grows to split it from above. It has been suggested that this splitting results from a Taylor instability [23] wherein a heavy fluid overlies a lighter one [24]. In applying the Taylor theory to bubbles in fluidized beds, the dense phase over the bubble is the heavy fluid and the gas in the bubble is the lighter fluid. An initial, random, small disturbance with amplitude η 0 is assumed to perturb the essentially horizontal upper surface of the bubble and to grow according to η ⫽ η 0 exp(nt)

(12)


41

where n is a parameter depending on the physical properties of the system and on the wavelength of the disturbance. In gas fluidized beds, n may be obtained from [25] R 4 ⫹ 2R 2 ⫺ 4R ⫹ 1 ⫺

g ⫽0 (k 3ν 2p)

(13)

where k⫽

2π λ

(14)

is the wave number of the disturbance and R ⫽ 1 ⫹ n(ν p k 2)

(15)

A numerical solution for Eq. (13) has been presented elsewhere [25]. For each ν p, there is a specific wavelength called the ‘‘most dangerous’’ wavelength, λ max, which exhibits the maximum growth rate. Clift et al. [25] have noted that disturbances initiated on the roof of a bubble are swept around the periphery. In practice, a bubble does not split unless the disturbance has grown sufficiently before the tip of the growing spike reaches the side of the bubble. The likelihood of splitting can be estimated by comparing the time constant for the growth of a disturbance (i.e., t e ⫽ 1/n) with the maximum time available for growth [25]: t am ⫽

冢冣冢 R0 g

0.5

ln

1 tan(θ/2)

冣

(16)

If the time available for growth, t am, is greater than the required growth time, t e, the bubble is liable to split. Otherwise, the disturbance grows so slowly that it does not achieve an amplitude large enough to cause splitting before it is swept around the bubble equator. Observations of splitting bubbles suggest that disturbances usually develop in a regular pattern on either side of the node. Assuming a node is located λ/4 from the bubble nose so that the node is an antinode in the initial disturbance, θ⫽

λ π ⫽ 2θ kR 0

(17)

The relationship between the stable bubble frontal diameter, d f , and the corresponding ν p is shown in Fig. 2. It is clear from Eq. (13) and from Fig. 2 that ν p is the dominant factor determining the growth of the instabilities, the most dangerous wavelength, λ max, and, hence, the maximum stable bubble size, d bmax. Thus, the prediction of the effect of system properties on the bubble stability depends on the prediction of the effect of system properties on ν p. Smaller particles and higher pressures are known to lower ν p [26] and thus, according to the Taylor instability theory, result in smaller d f . The major difficulty in testing the ability of using the

42


FIG. 2 Effects of kinematic viscosity on stable bubble frontal diameters.

Taylor instability theory to explain observed limited bubble growth in high-pressure fluidized beds is the lack of data, correlations, and theory relating the kinematic viscosity of the dense phase, ν p, to system parameters. The relationship between the stable bubble frontal diameter, d f , and the corresponding ν p is shown in Fig. 2. The actual bubble volume is approximately 75% of that of a sphere of diameter equivalent to d f [22]. Hence, d f and d b are related: d f ⫽ 1.1d b

(18)

For Taylor instabilities governing the maximum stable bubble size in fluidized beds, d bmax is dependent on ν p. Because ν p decreases with an increase in ε D and ε D increases with an increase in pressure and temperature, particularly for fineparticle systems, increased pressure and temperature result in a decrease in d bmax. However, the effect of particle size must be taken into consideration because the effect of pressure and temperature on d bmax is minimal for coarse-particle systems compared to fine-particle systems. A suitable relationship for estimating µ D from variations in ε D for gas fluidized beds is needed in order to predict d bmax a priori. Although such a relationship is not currently available, calculations using equations developed for systems of solid spheres in liquids have been shown to be consistent with µ D needed to limit bubble growth in high-pressure systems [14]. Hence, no suitable relationship exists today to allow the calculation of d bmax for fine-particle systems as impacted by temperature and pressure. Nonetheless, experimentally determined values for d b indicate


43

that higher pressures and temperatures yield smaller-diameter bubbles in gas fluidized beds, particularly for fine-particle systems. It has generally been reported that the bubble frequency, f, increases with increases in temperature [27–29] and pressure [30]. These results are consistent with decreases in bubble size for increases in temperature and pressure.

Coarse-Particle Systems A generalized correlation for bubble size has been developed by Cai et al. [31] for pressurized fluidized-bed combustors (PFBC) on the basis of a comprehensive analysis of previous work. The correlation takes into account the different flow regimes at different pressures and gas velocities, as well as the special variation of bubble size within the lower-pressure range of the bubbling regime. For PFBC (P in Pa), d b ⫽ 0.1905h0.8P 0.06(u 0 ⫺ u mf)0.42 ⫻ {exp[⫺1.4 ⫻ 10⫺14P 2 ⫺ 0.25(u 0 ⫺ u mf)2 ⫺ 1 ⫻ 10⫺6P(u 0 ⫺ u mf)]} (19) The equivalent bubble size of the entire bed is 0.06 (u 0 ⫺ u mf)0.42 d be ⫽ 0.1002L 0.8 f P

⫻ {exp[⫺1.4 ⫻ 10⫺14P 2 ⫺ 0.25(u 0 ⫺ u mf)2 ⫺ 1 ⫻ 10⫺6P(u 0 ⫺ u mf)]} (20)

Particle Entrainment and Carryover Entrainment occurs when gas bubbles break at the top of the fluidized bed and throw particles up into the freeboard region above the bed surface. At low gas velocities, these particles quickly fall back into the bed and are retained, but as the fluidizing velocity is increased, more and more particles are transported to greater heights above the bed surface and there exists a particle density gradient extending some distance above the surface. For sufficiently tall freeboards, there will be a certain height at which the density gradient eventually falls to zero, and above this height, the entrainment flux will be constant. This height is called the transport disengagement height (TDH). If the bed solids have a wide size distribution and the gas velocity in the freeboard exceeds the terminal fall velocity of the smaller ones, then these will be carried out of the system or elutriated. Entrainment from fluidized beds is affected by changing pressure and temperature. An increase in the operating pressure increases the carrying capacity of the gas and increases the amount of solids carried over [32,33]. Chan and Knowlton [32] also found that the TDH increased linearly with pressure in the range 1–30 atm. Findlay and Knowlton [20] studied the effect of increasing temperature on entrainment. They investigated temperatures in the range of ambient to 1033 K

44

Pressure and Temperature in Fluidized Beds while maintaining the system pressure constant; thus, the dominant effect was the increase in gas viscosity. The increased gas viscosity decreased the terminal fall velocity and increased the rate of entrainment for a given fluidizing velocity. For the discharge point exceeding TDH, Geldart [11] recommended various elutriation rate constants for particle size fraction d pi. The Zenz–Weil [34] equation is recommended for particles of diameter d p ⬍ 100 µm and superficial gas velocities u 0 ⬍ 1.2 m/s when the entire bed is potentially entrainable (even the largest particles have a terminal velocity u t ⬍ u 0). Some of the applicable materials include cracking catalyst, coal char, and other low-density solids:

冢冣

u 20 K*i ∞ ⫽A (µ gu 0) gd pi ρ 2p

B

(21)

where A ⫽ 1.26 ⫻ 107 and B ⫽ 1.88 when [(u 20 /(gd pi ρ 2ρ) ⬍ 3 ⫻ 10⫺4 A ⫽ 4.31 ⫻ 104 and B ⫽ 1.18 when [(u 20 /(gd pi ρ 2ρ) ⬎ 3 ⫻ 10⫺4 Geldart [11] proposed that

冢冣

K*i ∞ ⫺B′u t ⫽ A′ exp (ρ gu 0) u0

(22)

where A′ ⫽ 23.7 and B′ ⫽ 5.4 for higher velocities and coarser particles A′ ⫽ 31.4 and B′ ⫽ 4.27 for beds consisting largely of 1-mm solids A′ ⫽ 49.1 and B′ ⫽ 4 for 2.5-mm coarse particles in the bed It is clear that the elutriation rate constants will increase for increases in gas viscosity and gas density according to Eqs. (21) and (22). Also, the elutriation rate will increase for a decrease in the particle terminal velocity, u t, according to Eq. (22). Increasing the temperature will cause a reduction in the terminal velocity for particles in the Stokes region (small particles) and an increase for those in Newton region (large particles). Terminal velocities are unaffected by the gas density in the Stokes region (small particles), slightly in the transition region, and in the Newton region (large particles) according to (1/ρ g)0.5. Hence, as the gas pressure is increased at constant gas velocity, the size of the particles carried over will increase slightly.

Effect of Hydrodynamics on Reaction Conversion Because the conversion of fluidized-bed chemical reactors depends on the contact of solids (either catalyst or reacting) with reactant gases, it is advantageous to be


45

able to improve the amount of feed gas contacting the solids. This contacting can be improved by forcing more feed gas to flow directly into the dense phase (i.e., increasing q D relative to q B) and to keep the bubbles, which do form, small (improved mass transfer between dilute and dense phases). Because higher pressures increase gas–solid contacting for fine-particle systems, one would expect that higher pressures would result in improved reactor conversions. Weimer et al. [21] presented the only comparative results ever reported for a pilot-scale tubular fixed-bed (3.8 cm inner diameter ⫻ 22 cm long) and fluidizedbed (15 cm inner diameter ⫻ 2.7 m long) catalytic process operating at comparative gas hourly space velocities (GHSV) and temperatures. They investigated the exothermic Fischer–Tropsch synthesis of syngas at pressures of P ⫽ 3549 and 6996 kPa and temperatures between T ⫽ 613 and 663 K. The fluidizable particle size was measured to be around d p ⫽ 100 µm. Their results showed that for nominal pressures of P ⫽ 3549 kPa, the conversion of carbon monoxide (X CO) was noticeably higher for the fixed-tube reactor. For GHSV ⬃1000 h⫺1 and T ⫽ 623 K, X CO ⫽ 0.7 in the pilot fixed bed versus X CO ⫽ 0.55 for the pilot fluidized bed. At nominal P ⫽ 6996 kPa, the fixed-bed conversions were only slightly higher than those in the fluidized bed were. For GHSV ⬃1600 h⫺1 and T ⫽ 621 K, X CO ⫽ 0.74 in the fixed bed versus X CO ⫽ 0.71 in the fluidized bed. The better agreement between results at nominal P ⫽ 6996 kPa versus P ⫽ 3549 kPa was believed due to smaller bubbles at the higher pressures. Bubbles were reported to be d b ⬃5 cm in diameter for the P ⫽ 3549 kPa operation and to be d b ⬍ 2 cm in diameter for the P ⫽ 6996 kPa operation. The number of gas interchanges for bubbles traversing the bed (i.e., the number of times the bubble gas was exchanged with the dense phase gas as bubbles rise in the bed) was calculated to be between 5 and 6 for the smaller bubbles at higher pressures versus 1 for the larger bubbles at lower pressures. The investigators believed that even better results could have been achieved if finer particles were being fluidized (⬍100 µm). These results indicate that the smaller bubbles resulting from higher pressures in fine-particle systems contribute to substantial improvement in mass interchange, allowing fixed-bed-type conversions. The scale-up of high-pressure, fine-particle fluidized beds from laboratory data is simplified because bubble growth is limited by hydrodynamics. Normally, fluidized-bed scale-up is complicated by the fact that the wall for small-particle systems limits bubble growth and that bubbles grow as the bed gets larger. Results reported in the literature indicate that the highpressure, fluidized-bed bubble size can be controlled by proper selection of fineparticle size.

Nomenclature A A′ Ar

Constant in Eq. (21) Constant in Eq. (22) Archimedes number ⫽ d 3pρ g(ρ p ⫺ ρ g)g/µ 2g

46

Pressure and Temperature in Fluidized Beds B B′ db C1 C2 d bmax df do d be dp d eff d pi F Fr f g h hj K*i ∞ k Lf n P qB qD q G(in) R R0 Re mf T T (amb) t am te u0 u mf ut VB Vj X CO

Constant in Eq. (21) Constant in Eq. (22) Bubble diameter (m) C 1 ⫽ 42.86(1 ⫺ ε mf)/φ C 2 ⫽ φε 3mf /1.75 Maximum bubble diameter (m) Frontal bubble diameter (m) Jet nozzle or distributor orifice diameter (m) Equivalent bubble diameter, averaged over bed height (m) Particle diameter (m) Effective particle diameter based on experimental u mf calculation (m) Diameter of particle size fraction I (m) wt fraction of fine particles ⬍45 µm in diameter Froude number ⫽ u 20 /gd p Bubble frequency (1/s) Gravitational force (m/s2) Distance up the bed axially (m) Distance the jets protrude up the bed (m) Elutriation rate constant for size fraction d pi(kg/m2 s) Linear slope for rate of change of ε mf with temperature (l/K) Expanded bed height (m) Parameter used in Eq. (12) Pressure (Pa) Volumetric flow rate of gas as bubbles (m3/s) Volumetric flow rate of dense phase gas (m3/s) Volumetric flow rate of gas (m3/s) Defined by Eq. (15) Frontal radius of a bubble (m) Particle Reynolds number ⫽ d pu mf ρ g/µ g Temperature (K) Ambient temperature (K) Time available for growth (s) Required growth time (s) Superficial gas velocity (m/s) Minimum fluidization velocity (m/s) Terminal velocity (m/s) Bubble phase volume (m3) Jet phase volume (m3) Conversion of carbon monoxide

Greek εD ε mb ε mf ε mf(amb)

Dense phase voidage Voidage at minimum bubbling Voidage at minimum fluidization Voidage at minimum fluidization at ambient temperature

Pressure and Temperature in Fluidized Beds η η0 φ λ λ max µg νp θ ρg ρp

47

Amplitude of disturbance for Taylor theory [Eq. (12)] Initial amplitude of disturbance for Taylor theory [Eq. (12)] Particle sphericity Wavelength of disturbance for Taylor theory [Eq. (12)] Maximum wavelength of disturbance Gas viscosity (kg/m s) Kinematic dense phase viscosity Angular position where disturbance originates Gas density (kg/m3) Particle density (kg/m3)

References

1. T. M. Knowlton, ‘‘Pressure and Temperature Effects in Fluid-Particle Systems,’’ in Fluidization VII (O. E. Potter and D. J. Nicklin, eds.), Engineering Foundation, New York, 1992, pp. 27–46. 2. J. G. Yates, ‘‘Effects of Temperature and Pressure on Gas–Solid Fluidization,’’ Chem. Eng. Sci., 51(2), 167–205 (1996). 3. A. Mathur, S. C. Saxena, and Z. F. Zhang, ‘‘Hydrodynamic Characteristics of Gas– Solid Fluidized Beds over a Broad Temperature Range,’’ Powder Technol., 47, 247– 256 (1986). 4. C. Y. Wen and Y. H. Yu, ‘‘Mechanics of Fluidization,’’ Chem. Eng. Progr. Symp. Ser., 62, 100–111 (1966). 5. D. C. Chichester, R. M. Kornosky, L.-S. Fan, and J. P. Danko, ‘‘Characteristics of Fluidization at High Pressure,’’ Chem. Eng. Sci., 39(2), 253–261 (1984). 6. P. A. Olowson and A. E. Almstedt, ‘‘Influence of Pressure on the Minimum Fluidization Velocity,’’ Chem. Eng. Sci., 46(2), 637–640 (1991). 7. J. S. M. Botterill, Y. Toman, and K. R. Yuregir, ‘‘The Effect of Operating Temperature on the Velocity of Minimum Fluidization Bed Voidage and General Behaviour,’’ Powder Technol., 31, 101–110 (1982). 8. B. Formisani, R. Girimonte, and L. Mancuso, ‘‘Analysis of the Fluidization Process of Particle Beds at High Temperature,’’ Chem. Eng. Sci., 53(5), 951–961 (1998). 9. A. Lucas, J. Arnaldos, J. Casal, and L. Puigjaner, ‘‘High Temperature Incipient Fluidization in Mono and Polydisperse Systems,’’ Chem. Eng. Commun., 41, 121–132 (1986). 10. G. Raso, M. D’Amore, B. Formisani, and P. G. Lignola, ‘‘The Influence of Temperature on the Properties of the Particulate Phase at Incipient Fluidization,’’ Powder Technol., 72, 71–76 (1992). 11. D. Geldart, ‘‘Particle Entrainment and Carryover,’’ in Gas Fluidization Technology (D. Geldart, ed.), Wiley, Chichester, 1986, pp. 123–154. 12. A. R. Abrahamsen and D. Geldart, ‘‘Behavior of Gas Fluidized Beds of Fine Powders I. Homogeneous Expansion,’’ Powder Technol., 26, 35–46 (1980). 13. K. V. Jacob and A. W. Weimer, ‘‘High-Pressure Particulate Expansion and Minimum Bubbling of Fine Carbon Powders,’’ AIChE J., 33(10), 1698–1707 (1987). 14. A. W. Weimer and G. J. Quarderer, ‘‘On Dense Phase Voidage and Bubble Size in High Pressure Fluidized Beds of Fine Powders,’’ AIChE J., 31(6), 1019–1028 (1985).

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Pressure and Temperature in Fluidized Beds 15. A. Kmiec, ‘‘Equilibrium of Forces in a Fluidized Bed—Experimental Verification,’’ Chem. Eng. J., 23, 133 (1982). 16. L. A. Behie and P. Kehoe, ‘‘The Grid Region in a Fluidized Bed Reactor,’’ AIChE J., 19, 1070 (1973). 17. L. A. Behie, M. A. Bergougnou, and C. G. J. Baker, ‘‘Mass Transfer from a Grid Jet in a Large Gas Fluidized Bed,’’ in Fluidization Technology, Vol. 1, D. L. Keairns, ed.), Hemisphere, Washington, DC, 1976, p. 261. 18. I. Hirsan, C. Shishtla, and T. M. Knowlton, ‘‘The Effect of Bed and Jet Parameters on Vertical Jet Penetration Length in Gas Fluidized Beds,’’ 73rd Annual AIChE Meeting, Chicago, IL, 1980. 19. J. G. Yates, V. Bejcek, and D. J. Cheesman, ‘‘Jet Penetration into Fluidized Beds at Elevated Pressures, Fluidization 5 (K. Ostergaard and S. Sorenson, eds.), Engineering Foundation, New York, 1986, pp. 79–86. 20. J. G. Findlay and T. M. Knowlton, Final Report for Department of Energy, Project DE-AC21-83MC20314 (1985). 21. A. W. Weimer, G. J. Quarderer, G. A. Cochran, and M. M. Conway, ‘‘Design and Performance of a High-Pressure Fischer-Tropsch Fluidized Bed Reactor,’’ in Fluidization VI (J. R. Grace, L. W. Shemilt, and M. A. Bergougnou, eds.), Engineering Foundation, New York, 1989. 22. P. N. Rowe and B. A. Partridge, ‘‘An X-ray Study of Bubbles in Fluidized Beds,’’ Trans. Inst. Chem. Eng., 43, 157 (1965). 23. G. I. Taylor, ‘‘The Instability of Liquid Surfaces When Accelerated in a Direction Perpendicular to Their Planes,’’ Proc. Roy. Soc., A201, 192 (1950). 24. R. Clift and J. R. Grace, ‘‘The Mechanism of Bubble Break-up in Fluidized Beds,’’ Chem. Eng. Sci., 27, 2309 (1972). 25. R. Clift, J. R. Grace, and M. E. Weber, ‘‘Stability of Bubbles in Fluidized Beds,’’ Ind. Chem. Fundam., 13, 45 (1974). 26. D. F. King, F. R. G. Mitchell, and D. Harrison, ‘‘Dense Phase Viscosities of Fluidized Beds at Elevated Pressures,’’ Powder Technol., 28, 55 (1981). 27. T. Mii, K. Yoshida, and D. Kunii, ‘‘Temperature Effects on the Characteristics of Fluidized Beds,’’ J. Chem. Eng. Japan, 6, 100–102 (1973). 28. T. Otake, S. Tone, M. Kawashima, and T. Shibata, ‘‘Behaviour of Rising Bubbles in a Gas-Fluidized Bed at Elevated Temperature,’’ J. Chem. Eng. Japan, 8, 388– 392. 29. K. Yoshida, T. Ueno, and D. Kunii, ‘‘Mechanism of Bed to Wall Heat Transfer in a Fluidized Bed at High Temperatures,’’ Chem. Eng. Sci., 29, 77–82 (1974). 30. I. H. Chan, C. Shishtla, and T. M. Knowlton, ‘‘The Effect of Pressure on Bubbling Gas Fluidized Beds,’’ Powder Technol., 53, 217–235 (1987). 31. P. Cai, M. Schiavetti, G. De Michele, G. C. Grazzini, and M. Miccii, ‘‘Quantitative Estimation of Bubble Size in PFBC,’’ Powder Technol., 80, 99–109 (1994). 32. I. H. Chan and T. M. Knowlton, ‘‘The Effect of Pressure on Entrainment from Bubbling Gas Fluidized Beds,’’ in Fluidization (D. Kunii and R. Toei, eds.), Engineering Foundation, New York, 1984, pp. 283–290. 33. S. T. Pemberton and J. F. Davidson, ‘‘Elutriation of Fine Particles from Bubbling Fluidized Beds,’’ in Fluidization (D. Kunii and R. Toei, eds.), Engineering Foundation, New York, 1983, pp. 275–282. 34. F. A. Zenz and N. A. Weil, AIChE J., 4, 472 (1958). ALAN W. WEIMER

Process Safety and Risk Management

49

Fundamentals of Process Safety and Risk Management

Introduction The growth of industry and, as a result, the economy are dependent on technology advances and innovations. However, these same activities often lead to more complex processes, especially in the chemical industry, which is using comparatively severe operating conditions (temperature, pressure, flow rate, etc.), more reactive chemicals, and exotic chemistry. These more complex processes require in-depth analysis and knowledge of process chemistry and hazards. It is even more important now to design the process and equipment to precise standards based on a complete understanding of the underlying hazards, process chemistry, and the impact of operating conditions. Recently, much attention has been paid to human factors and its impact on chemical plant incidents. However, one can also say that process knowledge and understanding is the most human factor. This is based on the concept that inadequate knowledge, information, and understanding of the process hazards, chemistry, and impact of operating conditions are the root cause of many process plant incidents. Managing safety is no easy task, but it makes bottom-line sense. There is a direct payoff in savings on a company’s workers’ compensation insurance, whose premiums are partly based on the number of claims paid for job injuries [1]. The indirect benefits are far larger, for safe plants tend to be well run in general and more productive. The recipe for safety is remarkably consistent from industry to industry. It starts with sustained support of top management followed by implementation of appropriate programs and practices that institutionalize safety as a culture as compared to add-on procedures. The ingraining of safety as second nature in day-to-day activities requires a paradigm shift and can only be accomplished when safety is viewed as an integral and comprehensive part of any activity as compared to being a stand-alone or add-on activity.

Accident Process and Multiple Barrier Concept Most chemical plant accidents follow a typical pattern. It is important to study these patterns in order to be able to develop management systems to prevent these accidents. Also, many accidents occur as a result of the failure of multiple systems or ‘‘barriers.’’ In fact, it can be argued that many of these accidents may not have occurred, had at least one of the ‘‘barriers’’ not failed. Thus, it is important to study the concept of multiple barriers and its role in preventing process plant accidents.

50

Process Safety and Risk Management The Accident Process Most chemical accidents follow a three-step process, as described by Crowl and Louvar [2]:

• • •

Initiation: the event, which starts the accident process Propagation: the event, series of events, or condition which allows the accident process to continue, or which expands the magnitude of the accident Termination: the event or events, which stop the accident

The following is an example of the process:

• • • • •

A seal on a sulfuric acid pump leaked, requiring replacement (initiating event). The pump was drained and washed, but some time passed before maintenance began (propagating event). An isolation valve between the pump and the sulfuric acid supply was leaking (propagating event). The mechanic wore most of the required personal protective equipment, but failed to wear rubber boots (propagating event). When the mechanic began to work on the pump, he was splashed on the foot when a small amount of sulfuric acid was released, resulting in an acid burn (terminating event—all of the acid in the pump was released).

To prevent accidents, we must modify this accident process. This can be done by eliminating or reducing the likelihood of initiating events or propagating events, reducing the ability of propagating events to increase the magnitude of the accident, or by providing terminating events to interrupt the accident sequence before unacceptable consequences can occur. For the example described, some corrective actions might include the following:

• • •

Using a pump with an improved design, which would require less frequent seal repair (reducing the likelihood of the initiating event) Providing a double block between the sulfuric acid supply and the pump, and improving procedures and training to ensure timely washing of equipment and use of protective equipment (reducing likelihood of propagating events) Training the mechanic to assume the pump contains sulfuric acid and to drain it to a safe place before he begins his work (provide a safe terminating event by safely removing the acid)

Multiple Barrier Concept (Layers of Protection) Chemical processes traditionally rely on multiple layers of protections, or barriers, between a hazardous agent and the people, environment, and property which might be adversely impacted by an incident. This concept is illustrated in Fig. 1


51

FIG. 1 Typical layers of protection for a chemical process. (Based on Fig. 2.2 of Ref. 3.)

[3]. The layers of protection might include the basic process design, basic process controls and operating procedures, critical alarms and process shutdown procedures, safety interlocks, emergency equipment such as rupture disks and pressure relief valves, physical containment systems such as catch tanks and spill containment dikes, emergency response equipment and services such as sprinkler sys-

52

Process Safety and Risk Management tems and fire-fighting equipment and personnel, and personnel evacuation procedures. Multiple barriers are generally required because no barrier will be perfect— all are subject to potential failure. An inherently safer process (discussed elsewhere in this article) will reduce or eliminate the hazard and will require fewer or less robust layers of protection—and, if the hazard is sufficiently small, there may be no need for additional protective layers at all. This is highly desirable because the layers of protection may require significant initial capital investment and ongoing operating costs to ensure their continued effectiveness. Also, although the layers of protection may be highly reliable and the risk of an accident may be small, it can never be zero—there is always a possibility that all of the layers of protection will fail simultaneously and the accident will occur. The number and required reliability of the barriers or layers of protection must be established through the use of the various hazard and risk analysis techniques described in the following sections. This requires a complete understanding of the hazards of the process and plant-hazard identification, and an understanding of the mechanisms or scenarios by which those hazards might result in harm to people, the environment, or property—hazard analysis or hazard evaluation.

Regulations During the past 15 years, a number of chemical or related incidents in the petrochemical industry have adversely affected surrounding communities. A few of these incidents, such as the vapor cloud explosion in Flixborough in 1974, the liquefied petroleum gas explosion in Mexico City in 1984, the toxic material release in Bhopal in 1984, and the fire and radiation release in Chernobyl, were reported worldwide. Both governmental agencies and trade organizations responded by developing standards and regulations to improve process safety. The American Petroleum Institute (API) and the American Chemistry Council (ACC) started to work with their members to develop organizational guidelines. The U.S. Department of Labor directed the Occupational Safety and Hazard Administration (OSHA) to develop federal standards for managing process safety. A consensus started to emerge in 1990. Although the language, application, and extent of each document differed, the contents and objectives were almost the same. The API published Recommended Practice 750: Management of Process Hazards [4] in January 1990. OSHA published the proposed federal process safety rule [5] in July 1990. In October 1990, the ACC published its Resource Guide for Implementing the Process Safety Management Code of Practices [6]. In addition, the Clean Air Act Amendments of 1990 directed OSHA and the Environmental Protection Agency (EPA) to develop process safety management regulations to protect workers and the environment. The final OSHA rule on Process Safety Management of Hazardous Chemicals (29 CFR 1910.119) was published in the Federal Register [7] on February 24, 1992. A matrix showing the relevance of OSHA Process Safety Management (PSM) elements to the Center for Chemical Process


53

TABLE 1 Summary Comparison of OSHA Elements with CCPS Elements CCPS 12 elements of chemical process safety management 1. Accountability: Objectives and Goals 2. Process Knowledge and Documentation 3. Capital Project Review and Design Procedures (for new and existing plants, expansions, and acquisitions) 4. Process Risk Management 5. Management of Change 6. Process and Equipment Integrity

7. Human Factors 8. Training and Performance

9. 10. 11. 12.

Incident Investigation Standards, Codes, and Laws Audits and Corrective Actions Enhancement of Process Safety Knowledge

Relevant paragraphs of OSHAs PSM rule

Process Safety Information § 1910.119 (d) Pre-Startup Safety Review § 1910.119 (i)

Mechanical Integrity § 1910.119 (j) Process Hazard Analysis § 1910.119 (e) Pre-Startup Safety Review § 1910.119 (i) Management of Change § 1910.119 (l) Process Hazard Analysis § 1910.119 (e) Operating Procedures § 1910.119 (f ) Mechanical Integrity § 1910.119 (j) Process Hazard Analysis § 1910.119 (e) Operating Procedures § 1910.119 (f ) Operating Procedures § 1910.119 (f ) Training § 1910.119 (g) Pre-Startup Safety Review § 1910.119 (i) Emergency Planning and Response § 1910.119 (n) Incident Investigation § 1910.119 (m) Compliance Audits § 1910.119 (o)

Safety’s (CCPS) chemical process safety management elements is given in Table 1. EPA published the Risk Management Program in June 1996. The international chemical and petroleum community has also been addressing process safety management through regulations and recommended practices. The Norwegian Petroleum Directorate issued rules [8] in 1981 requiring quantitative hazard analyses for offshore petroleum operations. In response to the 1976 chemical dioxin release in Seveso, Italy, a European Directive [9] (commonly called the Seveso Directive) on process safety management was issued in 1982. More recently, the British government has issued process safety management regulations [10] for North Sea petroleum operations, following the recommendations of the widely distributed Cullen Report, which investigated the 1985 Piper Alpha offshore platform tragedy. Outside of Europe, the World Bank [11] has provided process safety management guidance for third-world projects. Similarly, the International Labor Office in Geneva has issued hazard analysis recommendations [12].

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Process Safety and Risk Management The Process Safety Management Program The 14 elements of the OSHA Process Safety Management (PSM) regulation (29 CFR 1910.119) were published in the Federal Register on February 24, 1992 [7]. The objective of the regulation is to prevent or minimize the consequences of catastrophic releases of toxic, reactive, flammable, or explosive chemicals. The regulation requires a comprehensive management program: a holistic approach that integrates technologies, procedures, and management practices. The process safety management regulation applies to processes that involve certain specified chemicals at or above threshold quantities, processes that involve flammable liquids or gases on-site in one location, in quantities of 10,000 lbs, or more (subject to few exceptions), and processes that involve the manufacture of explosives and pyrotechnics. Hydrocarbon fuels, which may be excluded if used solely as a fuel, are included if the fuel is part of a process covered by this regulation. In addition, the regulation does not apply to retail facilities, oil or gas well drilling or servicing operations, or normally unoccupied remote facilities. The process safety management regulation requires a systems approach for managing safety. Segments of the hazardous chemicals industry have for sometime practiced some or all of the required programs. The promulgation of the regulation formalized the requirements and established a minimum criterion. This is both good and bad. The regulation now requires everyone to establish the management systems and apply the technologies needed to comply with the regulation. However, because of the same reason, there is a tendency to look for ‘‘paper compliance’’ as compared to making real improvements in safety programs and technologies.

The Risk Management Program In 1996, the EPA promulgated the regulation for Risk Management Programs for Chemical Accident Release Prevention (40 CFR 68). This federal regulation was mandated by section 112(r) of the Clean Air Act Amendments of 1990. The regulation requires regulated facilities to develop and implement appropriate risk management programs to minimize the frequency and severity of chemical plant accidents. In keeping with regulatory trends, EPA required a performance-based approach toward compliance with the risk management program regulation. The EPA regulation also requires regulated facilities to develop a Risk Management Plan (RMP). The RMP includes a description of the hazard assessment, prevention program, and the emergency response program. Facilities submit the RMP to the EPA and, subsequently, it is made available to governmental agencies, the state emergency response commission, the local emergency planning committees, and communicated to the public. The risk management program regulation defines the worst-case release as the release of the largest quantity of a regulated substance from a vessel or process line failure, including administrative controls and passive mitigation that limit the total quantity involved or release rate. For gases, the worst-case release scenario assumes the quantity is released in 10 min. For liquids, the scenario assumes an


55

instantaneous spill and that the release rate to the air is the volatilization rate from a pool 1 cm deep unless passive mitigation systems contain the substance in a smaller area. For flammables, the scenario assumes an instantaneous release and a vapor cloud explosion using a 10% yield factor. For alternative scenarios (note: EPA used the term alternative scenario as compared to the term more-likely scenario used earlier in the proposed regulation), facilities may take credit for both passive and active mitigation systems. Appendix A of the final regulation lists endpoints for toxic substances to be used in worst-case and alternative scenario assessment. The toxic endpoints are based on ERPG-2 (Emergency Response Planning Guidelines—Level 2) or level of concern data compiled by the EPA. The flammable endpoints represent vapor cloud explosion distances based on overpressure of 1 psi or radiant heat distances based on exposure to 5 kW/m2 for 40 s.

Hazard and Risk Hazard A hazard is a physical or chemical characteristic of a material or process which has the potential to cause harm to people, the environment, or property. A hazard can be a characteristic property of a material, it can be a result of the conditions of use of the material, or it can be the result of an interaction among two or more materials or sources of energy. Some examples of hazards include the following:

• • • • • •

Chlorine is a toxic gas. Gasoline is a flammable liquid. Sulfuric acid is corrosive. A cylinder containing compressed air contains significant potential energy from the pressurized gas. A mixture of a vinyl monomer and a peroxide initiator has significant potential chemical energy of reaction. A 600 psig steam pipe is at elevated temperature and also contains a lot of energy from the pressure of the steam.

These hazards cannot be changed; they are intrinsic to the material or its conditions of use. The only way to eliminate or reduce hazards is to change the material or conditions of use. Although it is generally preferable to eliminate or reduce hazards (see later discussion on inherent safety), this is not always possible. The properties of a material or system, which create a hazard, may be the same as the properties, which make the material or system useful. A highly reactive monomer, when polymerized under controlled conditions, will produce a valuable product. However, if the polymerization is not controlled, the result could be overpressurization of a reactor and a possible explosion. Therefore, it is often necessary to manage and control the hazards of a process and plant. To do this, you must first identify and understand the hazards—hazard identification. Then, you must understand how

56

Process Safety and Risk Management the harm to people, the environment, or property can be realized from the hazardous material or condition—hazard evaluation or hazard analysis.

Risk Risk is a measure of human injury, environmental damage, or property loss expressed in terms of both the likelihood of the incident and the magnitude of the injury, damage, or loss. Risk can be considered to be a function of the potential incident occurrence, incident consequence, and incident likelihood: Risk ⫽ f (incident, consequence, likelihood) There can be many different kinds of risk associated with a chemical process or plant; for example, safety risk to plant workers, health risk to workers, health risk to neighbors, risk of various kinds of environmental damage, risk of damage to the plant or other property, risk of producing product which does not meet specifications and cannot be sold, risk of loss of business due to a plant outage, business risk that the product cannot be sold, and others. All of these risks must be understood and managed to successfully operate a profitable plant and business over the long term. There are many different measures for each of the risks associated with a chemical manufacturing facility. For example, some measures of risk to employees in a plant include the following:

• • • • •

Average risk of fatality from a process accident to an employee in the plant Maximum risk of fatality from a process accident to the employee at greatest risk Average risk to a specific employee in a plant over the course of a normal working day as he does various specific jobs An estimate of the distribution of likelihood of accidents of various size (impacting one employee, two employees, three employees, etc.) Average risk of injury to an employee

The CCPS Guidelines for Chemical Process Quantitative Risk Analysis [13] describes many different measures of safety risk which might be used in understanding the risk of a chemical plant and also provides quantitative methodologies for calculation.

Hazard Identification and Hazard Evaluation Objective of Hazard Identification and Evaluation The objective of hazard identification is to fully understand the hazards of a chemical process, including the hazards associated with the following:


• • •

57

Materials; for example, toxicity, reactivity Process conditions; for example, high temperature, high pressure Potential interactions among one or more materials; for example, chemical reaction, decomposition, corrosion

A full understanding of all hazards of a process is the essential first step in eliminating, minimizing, or managing those hazards. The objective of hazard evaluation is the identification of specific mechanisms by which the potential harm associated with the process hazards can be realized. Hazard evaluation techniques can also include the identification of protective measures which have been incorporated into the process design to manage the hazards, qualitative assessment of the risk of the specific incident scenarios identified, and evaluation of the adequacy of existing protective features or recommendation for additional safeguards.

Hazard Identification Hazard identification is based on a complete knowledge of the properties of the materials being handled and the chemical and physical processes used. Hazards of materials can be identified from literature searches, publications such as Sax’s Dangerous Properties of Industrial Materials [14], and libraries of Material Safety Data Sheets. The best source of hazard information for raw materials is often the material supplier. Bretherick’s Handbook of Reactive Chemical Hazards [15] provides a comprehensive summary of reactive chemical hazard literature. Checklists are a good mechanism for identifying process hazards (checklists as a hazard evaluation technique will be discussed later). Two specific tools for hazard identification, which are particularly useful in understanding chemical reaction hazards, are discussed in more detail in the following sections.

Interaction Matrix The interaction matrix (Fig. 2) is intended to identify chemical reaction hazards among materials and energy sources in a chemical process. This tool is particularly useful early in the development of a new chemical process. To create an interaction matrix, list all of the materials, materials of construction, likely contaminants, potential sources of energy, process utilities (such as steam, water, nitrogen, compressed air, ethylene glycol coolant, and heat transfer oil), and other relevant parameters along each axis of the matrix. It is a good idea to also include ‘‘people’’ on one of the axes, to prompt questions about toxicity and other adverse impacts of materials on people. Then, ask what happens for each interaction where the matrix columns and rows intersect. The matrix should go beyond a simple yes– no answer, but rather should provide some detailed information on the nature of the interactions identified. Often an interaction matrix will generate more questions than answers, particularly early in development. In this case, it may be appropriate to recommend a literature search or laboratory experiments to understand potential interactions.

58


FIG. 2 Example interaction matrix.

Chemistry Hazard Analysis The Chemistry Hazard Analysis (CHA) is derived from the Hazard and Operability (HAZOP) study methodology. The thought process of a HAZOP study can be applied at any stage in process development. The CHA is a HAZOP applied to a chemical reaction, without the detailed plant design information required for a traditional HAZOP study. For the CHA, the chemist or engineer usually assumes that the deviation identified by the application of the guide word to the chemical reaction does occur for some reason, not developing specific causes, and investigates the consequences. If the consequences are known, the designer should determine if they represent a hazard which must be understood and managed as a part of the process development, and document this information for future action or reference. In many cases, early in process development, the consequences may not be known, and additional research or experiments may be needed.

Hazard Evaluation: Selection of Procedure A number of different hazard evaluation techniques are in common use in the chemical industry, as listed in Table 2. Some techniques, particularly those based on logic models, require more detailed plant design information and it may not be possible to apply them early in process development. Table 3 provides some

Process Safety and Risk Management TABLE 2

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Process Hazard Analysis Tools Commonly Used in the Chemical Process Industries

Category

Process Hazard Analysis Tool

Description and Comments

Brainstorming techniques

Safety review Preliminary hazard analysis What If Hazard and Operability Study (HAZOP) Failure Mode and Effect Analysis (FMEA)

Checklist techniques

Checklist What If/checklist

Risk-ranking techniques

Relative ranking

Logic model techniques

Fault-tree analysis Event-tree analysis

Relatively unstructured brainstorming techniques to identify hazards and potential accident scenarios. A structured analysis procedure that focuses brainstorming activities, including use of a specific set of guide words or knowledge and checklists of known equipment failure modes. Predefined checklists based on previous experience compare a design to specific standards or good practice. When combined with What If analysis, the checklists are used to prompt brainstorming activities. A general category that includes a large number of quantitative and semiquantitative techniques which use checklists or equations based on material properties, quantities, and handling conditions to numerically rank risk. Examples include the Dow Fire and Explosion Index and the Dow Chemical Exposure Index. Logic models which identify specific causes combinations of events which lead to a potential accident scenario. These techniques require much detailed design information and usually focus on analyzing a few specific accident scenarios in detail. These techniques can be quantified and are important tools in quantitative risk analysis

guidance on how the various techniques have commonly been applied through the life cycle of a chemical process.

What-If Analysis What-If Analysis is a brainstorming technique in which a team with expertise on the process asks ‘‘what-if ’’ questions about the process to identify potential hazards or incident scenarios. The Preliminary Hazard Analysis and Safety Review techniques are forms of What-If Analysis. In a What-If Analysis, a team of experts on the process and plant meet in a free brainstorming session to ask ‘‘what-if ’’ questions to identify what can go wrong. The technique is very flexible and can

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60 Process Safety and Risk Management


61

TABLE 4 Example of the Results of a What-If Analysis for a Batch Process Hazard or consequence

What-If...? 1. Reactant feed rate is too high?

2. Reactor temperature is too low?

Safeguards

Recommendations

1. Heat generation 1. High reactant 1. Establish prevenrate exceeds heat flow rate intertive maintenance removal capabillock shuts down and testing proity, increased temfeed. High reactor gram for flow perature, potentemperature inand temperature tial runaway terlock shuts interlocks. reaction. down feed. Rupture disk sized adequately to protect reactor for maximum feed rate with no cooling. 2. Reaction may 2. Low reactor tem- 2. Evaluate rupture stall, resulting in perature alarm disk size relative a buildup of rewarns operator. to potential poolactants. This pool ing of unreacted of unreacted matematerial. Based rial has signifion result, detercant potential enmine if additional ergy, possible protection is runaway reaction needed. if temperature is subsequently increased.

be applied at any stage in the process life cycle. The unstructured nature of WhatIf Analysis can be both an advantage and a disadvantage. With an experienced and knowledgeable team, the technique can be powerful. The discussion and interaction among team members in the meeting can enhance the identification of hazards. However, the unstructured nature may result in an incomplete analysis by an inexperienced team. Table 4 is an example of the results of a What-If Analysis. WhatIf is often combined with a checklist to ensure that a minimum set of ‘‘What-If’’ questions are covered by the review team.

Checklist Analysis A checklist is a list of items used to verify that a plant or process is designed and operated consistently with a predetermined set of good practices defined by the checklist. A checklist is often used to confirm that a plant complies with codes, standards, or regulations. Checklists vary from general lists of questions describing common chemical hazards and processing concerns to detailed lists of specific requirements of a standard. Many checklists are simple, requiring only ‘‘yes, no,

62

Process Safety and Risk Management not applicable’’ answers. The use of checklist analysis depends on the availability of suitable checklists for the plant being reviewed. Good checklists are most likely to be available for common types of installation, such as flammable-solvent storage facilities, and are unlikely to be available for unique process operations. Checklist completeness depends on the experience of the checklist authors. The output of a checklist analysis is a list of responses to the checklist questions, with areas of noncompliance highlighted and recommendations for bringing the facility into compliance. Combining a What-If Analysis with a checklist can be a very effective hazard identification and evaluation technique. The What-If allows creative brainstorming of a team to identify hazards, and the checklist ensures that the team considers a specified list of hazards based on prior experience, addressing the concern about completeness of the What-If analysis.

Hazard and Operability Study A Hazard and Operability Study (HAZOP) is a guide word hazard evaluation technique normally done by a review team. HAZOP begins with the premise that the process is safe if operated, as intended, and the team must agree that this is true. Incidents are assumed to result from deviation from intended operation. Guide words are used in conjunction with the process operating parameters to identify potential deviations, and the review team determines the consequences of those deviations. A Hazard and Operability Study is best applied when specific process and plant information is available (e.g., a detailed plant design or an operating plant). To do a HAZOP, the process is first divided into sections, or nodes, which are analyzed individually. A node might be a transfer line from one vessel to another, a piece of process equipment such as a reactor or heat exchanger, or a step in a batch process. The team states the intended operation of each process node, including values of the process parameters—the process intention. The team then applies guide words to the parameters to identify potential deviations from intended operation. The following are basic guide words:

• • • • • • •

No More Less Part Of Reverse As Well As Other

As an example, the guide word ‘‘less’’ can be combined with a specified reactor temperature intention to arrive at the deviation ‘‘less (lower) reactor temperature.’’ Once a deviation has been identified, the team determines as many potential causes of the deviation as possible. For the deviation ‘‘less reactor temperature,’’ causes might include a cooling water control valve stuck open, incorrect temperature set point, and others. The team determines the consequences of each deviation, cause


63

combination, and existing safeguards. It then qualitatively judges the effectiveness of the safeguards to determine if they are adequate. A semiquantitative risk-ranking system is often used to aid the team in evaluating the significance of the hazards identified. If the existing safeguards are judged to be inadequate, the team should recommend appropriate action to mitigate the potential hazard. The team continues to apply the guide words to each node until no additional deviations can be identified. These steps are repeated for each process node, until the entire process has been reviewed. Table 5 shows a part of the output of a typical HAZOP study.

Fault-Tree Analysis A fault tree is a logic model which identifies the multiple ways in which equipment and human failures in a system can combine to cause an undesired event (the ‘‘Top Event’’). The analyst begins with a specific undesired event and develops a model using Boolean ‘‘AND’’ and ‘‘OR’’ logic gates to identify the immediate causes. These immediate causes can then be further developed to identify their causes, also using Boolean logic gates. The development of the tree continues until it reaches a level of resolution judged to be adequate for understanding potential incident sequences and identifying system improvements. These events are not further developed and are called basic events. Normally, basic events are at the level of failure of individual plant components (e.g., a shutoff valve stuck in the open position, a pump failure to run, a pressure sensor failing to detect high pressure). Human error can be incorporated into a fault tree by including specific operator actions (e.g., opening the wrong manual valve) or errors (e.g., failure to act in response to a high temperature alarm). Figure 3 is an example of the top levels of a fault tree for a fire. A fault tree can be solved using Boolean algebra techniques to identify specific combinations of individual equipment failures and human errors, which can cause the undesired top event. These combinations of failures and errors describe specific potential incident scenarios and are called minimal cut sets. The cut sets can be used to identify areas where the system can be improved. Fault trees can also be quantified by assigning failure rate and probability data to the basic events. These data can be mathematically manipulated to estimate the likelihood of the top event and to understand the relative contribution of individual basic events and cut sets to the total probability of failure.

Event-Tree Analysis An event tree is a graphical logic model which shows the possible outcomes resulting from an initiating event. An event tree describes the response of a system to a disturbance created by the initiating event. For example, Fig. 4 shows an event tree for the sulfuric acid splash incident used to describe the accident process. An event tree describes a number of potential outcomes from a single initiating event. These outcomes may vary in severity, and the event tree is useful in understanding the full range of possible outcomes that can result from a single system failure. Event trees are very useful in understanding the effectiveness of

A. Reactor R-310 temperature increases, heat balance indicates there is no potential for runaway reaction, even for maximum possible flow rate; product will not meet specifications if R-310 temperature increases above 90°C. B. Same as A. C. Same as A.

Consequences

A. 1. High flow alarm on FIC-301-01 warns operators. 2. High temperature alarm TAH-310-05 on Reactor R-310 warns operator. B. Same as A. C. A-2 only applies. A-1 will not provide an alarm because the flow sensor is miscalibrated. Plant experience indicates this is not likely.

Safeguards

A. Confirm that operator training includes proper response to FIC-301-01 and TAH-310-05 high alarms.

Recommendations

Note.—Intention: Feed 150–160 lbs./h of 42–44% aqueous Raw Material A solution from Feed Tank F-101 using Pump P-110 through line 10436 and flow controller FIC301-01 to Reactor R-310. Raw Material A solution temperature is 20–35°C. While feeding, Reactor R-310 agitator is running at 50 rpm.

A. Wrong set point (too high) on FIC-301-01 B. FIC-301-01 control value stuck open C. FIC-301-01 flow sensor miscalibrated—reads low

Causes

Example of Partial Results of a HAZOP

1. MORE than 160 lbs/h Raw Material A flow

Deviation

TABLE 5

64 Process Safety and Risk Management


65

FIG. 3 Top levels of a fault tree for a fire.

the multiple layers of protection, which are often present in a chemical process. Each independent layer of protection is a branch point in the event tree, with the branches corresponding to the success or failure of the layer of protection. An event tree, like a fault tree, can also be quantified by estimating a frequency for the initiating event and the probabilities of success and failure at each branch point in the event tree. Quantitative event tree analysis is often combined with fault-tree analysis: Fault trees are used to quantify the frequency of the initiating event and the probability of failure of the protective systems at the event-tree branch points.

Failure Mode and Effect Analysis A Failure Mode and Effect Analysis (FMEA) lists the known failure modes of specific pieces of equipment in a plant and determines the impact of those failures on the plant. FMEA and HAZOP are similar; the main difference is the starting point for identifying potential hazardous incident scenarios. HAZOP starts by postulating a deviation in the value of a process parameter (e.g., more flow) and asking what kind of equipment failures or operating errors might have caused that devia-

66


FIG. 4 Example event tree.

tion and what the process impact will be. FMEA starts by postulating a known equipment failure mode (e.g., control valve stuck open) and asks what impact this failure will have on the operation of the process. The FMEA starts with a functional description of each piece of process equipment and identifies ways in which that piece of equipment might fail to perform as designed. A good understanding of the equipment and potential failure modes is required. The FMEA determines how the process will respond to the potential equipment failure, determines if a potentially hazardous incident will result, identifies existing safeguards, evaluates their effectiveness, and develops recommendations for action where appropriate. These steps are very similar to the corresponding step in a HAZOP study. Table 6 shows a part of the output from a typical FMEA study.

Risk-Ranking Techniques Risk-ranking techniques such as the Dow Fire and Explosion Index [16] and the Dow Chemical Exposure Index [17] develop a numerical risk-ranking index based on various process characteristics such as material properties, chemical reactions, unit operations, operating conditions, and other factors. These risk indices provide a relative ranking of specific types of process hazards (e.g., fire and explosion hazards) and are useful for comparing alternate process or plant designs (including location and siting), understanding the parts of a plant or process which are the major sources of risk, and prioritizing other hazard evaluation and risk management activities.

1. Cooling Water Control Valve FIC-301-10. Pneumatically operated valve, fails closed on air supply failure A. Unable to increase flow of cooling water to reactor to control temperature of batch during exothermic reaction. Possible uncontrolled reaction generating high temperature and pressure. B. Valve will open, putting full cooling on the reactor.

B. Air failure to FIC-30110

Effects

A. Valve stuck in closed position

Failure mode

Example of Partial Results of a Failure Mode and Effects Analysis

Item and Description

TABLE 6

B. Full cooling is sufficient to control temperature at maximum raw material feed rate.

A. High temperature interlock TAH-310-08 on Reactor R-310 closes raw material feed valve.

Safeguards

A. Confirm that there is no buildup of unreacted material and stopping feed will stop reaction. Review maintenance procedures and make sure Valve FIC-301-10 is regularly inspected and tested.

Recommendations

Process Safety and Risk Management 67

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Process Safety and Risk Management Summary of Hazard Identification and Evaluation This section has briefly described a number of commonly used hazard identification and evaluation tools. Most of these tools are best used in a multidisciplinary team environment, providing a wide variety of plant and process experience and interactive discussion to understand the process and identify potential hazards and incident scenarios. CCPS [18] and Wells [19] provide more information on the application of these and other hazard evaluation tools in the chemical industry.

Consequence Analysis As mentioned earlier, risk is defined as a function of incident occurrence, frequency, and consequence. Consequence analysis is the quantitative estimation of the consequence of a chemical process incident—an estimate of the magnitude of the potential harm to people, the environment, or property. Because there is a wide range of potentially harmful impacts of chemical process incidents, there is a number of different tools which may be useful in analyzing these impacts. In this discussion, the consequence analysis tools described will be limited to those commonly used to estimate the potential for injury or fatality to people as an immediate result of exposure to harmful materials or energy. However, it is recognized that there is a wide variety of other potential consequences of incidents and a correspondingly wide variety of tools used to understand these consequences. Consequence models can be quite complex and can only be described in general terms in this discussion. A number of publications by the Center for Chemical Process Safety [20–24] describe specific types of consequence analysis models in detail. Les [25] also provides a detailed description of incident consequence models. There are also a number of public domain and proprietary commercial computer-modeling systems available for chemical release consequence analysis.

Source Models Source models quantitatively estimate the magnitude, rate, duration, physical state (solid, liquid, gas, or a combination), and temperature or other physical condition of a chemical release based on the physical and chemical parameters associated with a particular release scenario. Most source models are well developed in chemical engineering theory and are essentially the same as the models used for similar material flow scenarios used to design plant equipment. These include single-phase and multiphase flow models for flow-through holes, orifices, and pipes, which are readily adapted to describe flow from a leaking pipe or vessel. Two-phase flashing flow models are based on technology developed by the Design Institute for Emergency Relief Systems (DIERS) [26]. Two-phase or flashing jet release models must also consider the formation of fine aerosols in the discharge and the potential for the small drops


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to remain suspended in the atmosphere rather than ‘‘raining out’’ into an evaporating pool on the ground. For a discharge of material from a reactive system, the models required to fully understand the system may be quite complex and data from reaction calorimetry tests may be required. If a release is wholly or partially in the form of a liquid, it will form a pool on the ground. The evaporation of vapor from this pool is another potential source term for atmospheric dispersion models, which estimate the downwind concentration of the vapor. The first step in estimating evaporation from a liquid pool is to estimate the size of the pool. Pool size models consider the momentum of the liquid stream entering the pool, gravity spreading resulting from the depth of the pool, and the liquid physical properties (e.g., viscosity, surface tension, and surface wetting properties). Physical constraints such as dikes and containment systems may also determine the size of a pool of spilled liquid. There are three major pool evaporation situations, which are typically modeled: boiling liquid pools, volatile liquid pools, and relatively nonvolatile liquid pools.

•

• •

Boiling liquid pools occur when the pool liquid boils at a temperature below that of its surroundings (the ground and atmosphere). In this case, vapor generation is controlled by heat transfer into the liquid pool, both from the ground and from the surrounding atmosphere. The vapor release rate is determined from an estimate of the total heat transfer into the pool and the heat of vaporization of the liquid. Volatile liquid pools exert a significant vapor pressure but are at a temperature below the liquid boiling point. Evaporation models for volatile liquid pools consider both heat transfer into the pool and mass transfer rates into the atmosphere from the pool surface. The evaporation of relatively nonvolatile liquid pools is primarily determined by mass transfer at the surface of the pool. Because the evaporation rate is low, the pool temperature will be essentially the same as the temperature of the surroundings after any initial temperature differences equilibrate. Evaporation models are based on standard methodologies for estimating convective mass transfer from a liquid into a gas.

Vapor Cloud Dispersion Vapor cloud dispersion models estimate the area covered by the vapor cloud from a chemical release as it disperses in the atmosphere, and they estimate the vapor concentrations at specific locations in the cloud. Some of the data required for a vapor cloud dispersion model include the following:

• •

Characteristics of the release, including rate, total quantity released, location of the release Characteristics of the release (phase, direction, velocity, composition, temperature, pressure)

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• •

Atmospheric conditions, including wind speed, atmospheric stability, temperature, pressure Characteristics of the surface, including surface roughness

Some of the complex vapor cloud dispersion models may require additional information to characterize the release, the atmospheric conditions, and surface conditions. Vapor cloud dispersion models consider three typical types of behavior:

• • •

Neutrally buoyant gases (having a density close to the density of air) Positively buoyant gases (having a lower density than air) Dense or heavy (negatively buoyant) gases

Two major types of releases must also be considered:

• •

Instantaneous (puff releases) Continuous releases (plumes)

The CCPS [20] describes vapor cloud models in detail, including all of the major types of dispersion models and release types. The CCPS [22] provides a more condensed summary of some of these models. The output of these models describes the concentration of the released material in both time and space as the vapor cloud travels downwind.

Fires Incident consequence analysis may require consideration of one or more of several different types of fire:

• • •

Pool fire: a burning pool of a flammable or combustible liquid Jet fires: burning of a flowing jet of flammable liquid or gas, usually from a pipe or vessel Flash fire: nonexplosive combustion of a flammable mixture of a combustible vapor in air

Pool Fire The primary mechanism of damage from a pool fire is thermal radiation from the flame. Pool fire models estimate the thermal effects based on the properties of the material burning in the pool, the geometry of the pool, atmospheric characteristics, and geometry of the fire relative to the receiving source. Pool fire models are well developed. They are based on empirically determined characteristics such as burning rate, flame height, surface emissive power, and atmospheric transmissivity, all of which are well established in the literature. Pool fire models provide an estimate of the thermal radiation at locations of interest surrounding the fire.


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Jet Fire Many jet fire models are based on models used for the design of flare systems. As for pool fires, the damage from a jet fire results primarily from thermal radiation from the fire. Jet fire models require an understanding of the characteristics of the jet (discharge rate and velocity, material burning properties) and, like pool fires, characteristics of the atmosphere. Jet fire models are primarily empirical but are derived from much data and experience. The models produce an estimate of the thermal radiation at locations surrounding the jet fire.

Flash Fire A flash fire results from the ignition of a cloud of flammable gas (a cloud containing a flammable material at a concentration between its lower and upper explosive limits in air). Such a cloud can explode under the proper conditions (size and degree of confinement), resulting in a vapor cloud explosion. If the conditions required for a vapor cloud explosion are not present, the cloud may still ignite and burn. In this case, the burning cloud will not generate pressure and an explosion, but the flash fire is still capable of causing significant damage. The primary hazard is from direct contact with the flame and from thermal radiation, which is normally for a brief time of a few tenths of a second. Flash fires are normally modeled by determining the dimensions of the flammable cloud using vapor dispersion models and estimating the thermal radiation resulting from combustion of the cloud.

Explosions Chemical incident consequence analysis may need to consider four types of explosion:

• • • •

Physical explosions: the failure of a vessel containing material under pressure without chemical reaction (e.g., due to a vessel defect or excess pressure in the vessel). Vapor cloud explosions: explosion of a cloud of flammable vapor dispersed in the atmosphere. Confined explosions: explosion resulting from a rapid chemical reaction generating high temperature and pressure inside a confined space such as a vessel or a building. Boiling liquid expanding vapor explosions (BLEVE): the catastrophic failure of a vessel containing a superheated liquid. If the liquid is flammable, ignition may result in a fireball.

The CCPS [20,22] describes commonly used explosion models in detail.

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Physical Explosions Physical explosion models generally estimate the amount of energy which would be released by the sudden expansion of the material contained in a vessel from its initial temperature, pressure, and volume to atmospheric pressure. This estimated energy is then converted to an equivalent amount of TNT. A number of correlations of explosion pressure as a function of distance from a TNT explosion have been published, and these can be used to estimate damage. It may also be necessary to consider the potential impact of the vessel fragments, which result from a vessel explosion. Empirical models to estimate the number and size of fragments, their travel distance, and energy are available.

Vapor Cloud Explosions If ignited, a flammable vapor cloud can burn as a flash fire, or, if the flame speed accelerates sufficiently, it can produce significant blast pressure from a vapor cloud explosion. A number of factors have been found to be important in determining whether a vapor cloud explosion occurs when a flammable vapor cloud is ignited. These include the following:

• • • •

Turbulence in the vapor cloud. This turbulence may arise from the energy from the release of the fuel itself (from a jet or catastrophic loss of containment) or from the interaction of the cloud with its surroundings during the combustion process. Partial confinement of the vapor cloud as a result of obstacles, structures, or other factors, which could cause local partial confinement. The explosive combustion in the locally confined cloud can propagate into the rest of the cloud. Mass of the cloud. Experimental studies have demonstrated that there is a minimum mass of flammable material required to transition to a vapor cloud explosion. The CCPS [21] reports studies indicating that this minimum mass is in the range of 1 to 15 tons for typical hydrocarbons. Combustion properties of the fuel. Materials with a high fundamental burning velocity such as ethylene oxide and ethylene are reported to be more readily inclined to propagate to a vapor cloud explosion.

Vapor cloud explosions are modeled using three types of model:

•

TNT Equivalency Models. The total energy available from the combustion is estimated from the mass of fuel in the cloud and the heat of combustion of the fuel. This combustion energy is then converted to an equivalent mass of TNT and reduced by an ‘‘explosion efficiency’’ factor, which is empirically estimated. The explosion overpressure and other characteristics can then be estimated as a function of distance from the cloud using readily available experimental data for TNT explosions. TNT equivalency models are empirical, and the results are strongly dependent on the explosion efficiency, which may not be known for a particular material or cloud configuration. TNT equivalency models also do not characterize the vapor cloud explosion well in the area close


•

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to the cloud, where they may predict much higher pressure than typically result from the combustion of a flammable cloud. Multienergy Method. This model is based on the assumption that the blast characteristics of a flammable vapor cloud depend more on the level of congestion and confinement than on the fuel. The models require dispersion models to determine the size of the cloud. Then, areas with different confinement and congestion characteristics are identified and considered to be sources of strong blasts. The energy from each blast source is estimated, and the potential damage is estimated from empirically derived correlations. Baker–Strehlow Method. This model also considers confinement as the basis for the size of the flammable vapor cloud. It also considers burning characteristics and reactivity of the fuel, geometry of the confined volume, and the degree of confinement created by the obstacles in the confined volume. Blast characteristics are then estimated using a set of correlations and charts.

These models are discussed in detail in Refs. 20 and 22.

Confined Explosions Confined explosions result from combustion or another rapid chemical reaction in a confined vessel or building. Confined combustion reactions may occur with flammable vapor–air mixtures or from the dispersion of a cloud of combustible dust in air. The combustion or reaction products are often gases, and pressure is generated by the gas and also the elevated temperature resulting from the heat of combustion or reaction. Confined explosions are modeled by estimating the peak pressure that can be generated from the chemical reaction. The models are specific to the reaction and may require considerable reaction thermodynamic and kinetic data. The maximum pressure resulting from the reaction model is then compared to the failure characteristics of the confining vessel or building. If the pressure exceeds the expected failure pressure of the vessel, the damage resulting from vessel failure and the potential for damage or injury from fragments can be estimated using the methods for physical explosions discussed earlier.

Boiling Liquid Expanding Vapor Explosions A BLEVE is the rapid release of a large amount of superheated liquid to the atmosphere. It often occurs as a result of weakening of a pressure vessel caused by direct flame impingement on the vessel above the liquid level. This weakens the metal vessel and it can fail rapidly and catastrophically. The sudden loss of confinement allows the superheated liquid to rapidly flash, increasing its volume several hundred times and generating a pressure wave and fragments. If the released liquid is flammable, it can also ignite, resulting in a fireball. BLEVE models are based on the expansion energy of the flashing liquid. Blast effects tend to be local, and the impact of the fireball, which usually accompanies a BLEVE of a flammable material, is the more important source of damage. BLEVE fireball models empirically estimate the fireball dimensions based on the quantity of material released.

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Process Safety and Risk Management Thermal radiation characteristics of the fireball are then modeled using a combination of empirically derived relationships and fundamental models for the geometry of the fireball with respect to the receptor and atmospheric transmission of the thermal radiation. The result is an estimate of the radiant energy flux level and duration at various locations surrounding the BLEVE. Effect Models The result of the application of the models discussed in this section so far is an estimate of some type of physical parameter at various locations surrounding a chemical release: a concentration of toxic gas in the atmosphere, the amount of radiant energy at a specific location from a fire, and the peak pressure and impulse duration from an explosion. Effect models estimate the damage, which results from these physical effects. There are a wide range of possible effect models corresponding to the wide range of potential damage to people, the environment, and property, which can result from exposure to toxic materials, fires, and explosions. The CCPS [22] provides a summary of effect models commonly used to estimate the impact of toxic vapors, fires, and explosions on people. These models are generally empirical and are based on experimental data and evaluation of the consequences of past incidents. Models are available to estimate the impact of a hazardous agent using the dose-response relationship (e.g., relating probability of fatality to concentration and duration of exposure by inhalation of a toxic gas, relating severity of burns to intensity and duration of exposure to thermal radiation, or estimating damage to structures based on peak overpressure and duration).

Risk Assessment However many the resources we devote to the prevention of accidents, we can never eliminate every risk. We have to decide our priorities: Which risks should we deal with first? Which are so small compared with the other risks to which we are exposed that we should tolerate them, at least for the time being? Often, the judgment is qualitative: Some risks are illegal; some are obviously intolerable large or acceptably small; sometimes a generally accepted standard or code of practice tells us what to do. In other cases, the decision is not obvious and we use a numerical method known as quantitative risk assessment (QRA), probabilistic risk assessment (PRA), or, in the chemical industry, hazard analysis (Hazan). Stages of Risk Assessment Before carrying out a risk assessment, we have to identify the hazards (i.e., the substances, objects, or situations that can give rise to injury or damage) using one of the methods described earlier. (A risk, in contrast, is the probability that injury or damage will occur.) There are then three questions to answer:


• • •

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How often will injury or damage occur? What is the extent of the injury or damage? What action should we take?

Whenever possible, the answer to the first question should be based on experience, but often there is no experience, as the equipment is new or failure has never occurred. We then estimate a failure rate for the equipment as a whole, based on the known failure rates of its components, as described earlier. Similarly, the answer to the second question should be based on experience whenever possible but can be estimated as by one of the methods described earlier. The answer to the third question depends on the nature of the consequences. If damage is possible but injury is not, then the average cost of the damage (including consequential loss) is compared with the cost of prevention. If injury is possible, then the QRA approach is to set a target or criterion, usually based on the risk to life. Risks above a certain level should be removed or reduced as a matter of priority. Those below this level can be left alone, at least for the time being. Thus, QRA is a method for determining priorities. In a later development, there are two levels of risk. Risks above an upper level are considered intolerable; if they cannot be reduced, the plant should not be built (or should not be operated if it is already built). The risk considered tolerable for members of the public is much lower than that considered tolerable for employees. Risks below a much lower level are considered acceptable and need not be reduced. In between the two levels, we reduce the risks if we can, but we tolerate them if it is impracticable or very expensive to do so. The pressure to reduce them is great if the risk is near the intolerable level and reduces as we approach the acceptable level. The extent to which this approach is used and the risk levels are made explicit differs from country to country. The United Kingdom has long accepted the principle that we should compare the size of a risk with the cost, in money, time, and trouble, of removing it (although the ability to pay is not a deciding factor). If there is a gross disproportion between them, the risk being insignificant in relation to the cost, the risk can be tolerated. QRA was therefore accepted readily and the regulatory authority has suggested figures for the tolerable and acceptable risk levels. Other governments have been reluctant to admit that even trivial and infrequent risks should be tolerated and this has hindered the use of QRA. The actual risk levels suggested for the United Kingdom are as follows. They are similar to those used by many organizations elsewhere. Risk of death per person per year Maximum tolerable risk (employees) Maximum tolerable risk (public) Maximum tolerable risk (public—nuclear risks) Broadly acceptable risk (employees and public)

10⫺3 10⫺4 10⫺5 10⫺6

The maximum tolerable risk to employees seems rather high, but this risk is, in fact, tolerated in some industries.

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Process Safety and Risk Management For comparison, the annual risk of death from all causes is about 10⫺4 for someone aged 20 years and about 10⫺3 for someone aged 60 years.

Public Attitudes Quantitative risk assessment is difficult to explain to the public. They pick on the fact that a number of people could be killed in an industrial accident but ignore the fact that the probability that this will occur is extremely low. The death of 10 people once in 10 years is given far more publicity than the death of 1 person per year for 10 years. As a result, public pressure often compels industry and government to reduce risks which are already low but which the public perceives as high. At its best, this is democracy in action; at its worst, it is giving the most to those that shout the loudest. The public tends to oppose risks with the following traits:

• • • • • • • •

Imposed rather than accepted voluntarily Not under the individual’s control Of no obvious benefit to them Man-made rather than natural Unfamiliar Dreaded (e.g., cancer is more dreaded than heart disease though the latter kills far more people) Immoral (e.g., crime is feared more than road accidents) Associated with unpleasant events (e.g., nuclear power reminds us of atomic bombs)

When the public cannot judge the message, they judge the messenger. Unfortunately, most of these concerns make the man in the street oppose the chemical industry: The risks are imposed, not under his control, man-made, unfamiliar, and dreaded; past experience has been unpleasant; the industry does not obviously benefit him; and the spokesmen for the industry are often outsiders. There is no easy way of countering this perception. We try to explain the benefits of the industry and the low levels of risk, but we cannot say that accidents will never happen.

Incident Investigation The purpose of incident investigation is to find out why the incident occurred so that we can prevent it from happening again. The purpose is not to find out who should be blamed. Many people have an opportunity to prevent almost every incident. Figure 5 shows by example the opportunities that are available to prevent a fire or minimize the consequences of an apparently simple incident: An expansion joint (bellows) was incorrectly installed in a pipeline so that it was distorted. After some months, it leaked and a passing vehicle ignited the escaping vapor. Damage


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FIG. 5 An example of an accident chain. An expansion joint (bellows) was incorrectly installed so that it was distorted. After some months, it leaked and the escaping vapor was ignited by a passing vehicle. Damage was extensive, as the surrounding equipment had not been fireprotected to save the cost. Many people in various functions could have prevented the incident or minimized the consequences.

was extensive, as the surrounding equipment had not been fire-protected to save the cost. Many people could have prevented the fire, not just the fitter who installed the expansion joint incorrectly. The fire could have been prevented by better detailed design (not using expansion joints for hazardous materials), by better design methods (using HAZOP, consulting experts, better design standards, better training of designers), by better training of the fitter, by better inspection of workmanship, by keeping eyes open on plant visits, and by not tolerating poor workmanship in the past. We should investigate all incidents, including those, which, by good fortune, caused no injury or damage, but might easily have done so. Next time, they may.

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Process Safety and Risk Management Finding the Facts

• • • • • • •

Include people with a variety of experience on the investigating panel. It should not be too large; four or five people are usually sufficient. Do not disturb evidence that may be useful to experts who may be called in later. Draw up a list of everyone who may be able to help, such as witnesses, experts, designers, and people on other shifts. Be patient when questioning witnesses. Valuable information may be missed if we try to take police-type statements. Do not put ideas into people’s minds. Avoid questions to which the answer is ‘‘yes’’ or ‘‘no.’’ Make it clear that the objective of the investigation is to find out the facts, so that we can prevent the incident happening again, not to establish blame. Inform any authorities who have to be notified. Record information on damage and injuries so that others can use it for prediction.

Drawing Conclusions from the Facts Accident investigation is like peeling an onion. Beneath the immediate technical causes, look for ways of avoiding the hazard, such as inherently safer design. Look also for weaknesses in management, such as poor training or instructions or turning a blind eye to previous failures to follow instructions. Concentrate on prevention rather than causes. Look for causes that lead to actions. Do not, for example, quote corrosion as a cause and stop there. Ask if it was foreseen. If not, why? If it was foreseen, why did it occur? Was the right material of construction used? Were operating conditions outside the design range? Was monitoring carried out? If so, were the results followed up? Preventing leaks is a more effective way of preventing liquid and gas fires than removing sources of ignition (although we should also do what we can to remove known sources of ignition). Avoid the use of the term ‘‘human error’’ and never recommend someone to take more care. Instead, ask if we need better training, better instructions, or better compliance with instructions and, if so, say how this will be achieved. If an error was due to a slip or lapse of attention, inevitable from time to time, look for ways of removing opportunities for error. It is often useful, especially when investigating fires and explosions, to ask why it occurred when it did and not at some other time. It is also useful to ask if similar incidents, perhaps with less serious results, have occurred before, what recommendations were then made, and if they were effective or allowed to lapse. Avoid long shopping lists of possible recommendations. Ask if the cost of each recommendation is proportional to the size of the risk. Consider alternative solutions as well as the obvious ones. For each recommendation, make it clear who will carry it out and when. Bring the report forward at that time. Otherwise, nothing will happen except a repeat of the incident. Managers should not accept reports that fall short in any of these respects.


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They should look out for what is not said. For example, writers of accident reports are naturally reluctant to draw attention to similar incidents that had occurred elsewhere and, if they had been followed up, could have prevented the accident.

Spreading the Message Many companies restrict the circulation of incident reports, but this will not prevent the incident from happening again. We should circulate the essential messages throughout the company. There is no need to say where the incident occurred. Remember that incident reports grab people’s attention and are read, whereas advice and instruction are put aside to be read when we have time (if we ever do). Having paid the high price of an accident, we can recover some of the cost by turning it into a learning experience. Circulate reports containing new or forgotten information throughout the industry, so that others can learn from them. There are several reasons for doing so.

• • • •

Moral: If we have information that might prevent another accident, we have a duty to pass it on. Pragmatic: If we tell other organizations about our accidents, they may tell us about theirs. Economic: We would like our competitors to spend as much as we do on safety. The industry is one: Every accident affects its reputation.

Remembering the Message Incident reports are written, acted on, and then filed and forgotten. After a few years, people forget the reasons for the changes that were made. Procedures lapse or the equipment falls out of use and the incident happens again, even in the plant where it happened before. To prevent this from happening we should do the following:

• • • • • • •

Include in every instruction, code, and standard a note on the reasons for it and accounts of accidents that would not have occurred if the instruction, code, or standard had been followed. Never remove equipment before you know why it was installed. Never abandon a procedure before you know why it was adopted. Describe prior accidents as well as recent ones in safety bulletins and discuss them at safety meetings. Giving the message once is not enough. Follow up at regular intervals to see that the recommendations made after accidents are being followed, in design as well as operations. Remember that the first step down the road to an accident occurs when someone turns a blind eye to a missing blind. Include important accidents of the past in the training of undergraduates and company employees. Keep in every control room a folder of reports on past accidents. It should be read by all new arrivals and others should browse it during quiet shifts.

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Devise better retrieval systems so that we can find, more easily than at present, details of past accidents in our own and other companies and the recommendations made afterward.

The Management of Safety Inherently Safer Design The first step in the management of safety, after the hazards have been identified (see the section Hazard Identification and Hazard Evaluation), is to see if they can be removed. Only when we cannot do so, should we look for ways of keeping them under control or mitigating their consequences. When we remove a hazard, the safety is inherent in the design and cannot be lost. When we control a hazard, the protective equipment may fail, or be neglected, or the safety procedures may lapse. Note that we refer to inherently safer, not safe, design as we can rarely, if ever, remove every hazard. The principle routes to inherently safer design are as follows:

•

Intensification or minimization: Using so little hazardous material that it does not matter if it all leaks out. ‘‘What you don’t have, can’t leak.’’ This may seem obvious but until the explosion at Flixborough, UK in 1974 little thought was given to ways of reducing the amount of hazardous material in a plant. Engineers simply designed a plant and accepted whatever inventories the design required, confident that they could keep it under control. Flixborough weakened that confidence, and 10 years later, Bhopal almost destroyed it.

Microreactors promise much greater intensification than has been possible in the past. Intensification, when it is practicable, is the first choice, as it brings about greater reductions in cost. If less material is present, we need smaller pipes and vessels and smaller structures and foundations.

• •

•

Substitution: If intensification is not possible, then an alternative is substitution, using a safer material in place of a hazardous one. Thus, it may be possible to replace flammable solvents by nonflammable ones and processes that use hazardous raw materials or intermediates by processes that use safer ones. Attenuation or moderation: Another alternative to intensification is attenuation—using a hazardous material under the least hazardous conditions. Thus, liquefied chlorine and ammonia can be stored as refrigerated liquids at atmospheric pressure instead of storing them under pressure at ambient temperature. The lower pressure results in smaller leaks through a hole of a given size leak and the lower temperature results in less evaporation. Limitation of effects, by changing designs or reaction conditions rather than by adding on protective equipment, which may fail or be neglected. For example, it is better to prevent overheating by using steam or oil at a safe temperature than by using a hotter medium and a control system.


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Simplicity: Simpler plants are safer than complex plants, as they provide fewer opportunities for error and contain less equipment that develop faults. They are usually also cheaper.

Defense in Depth The hazards that cannot be removed have to be controlled. Because we depend on equipment and people, both of which may fail, we use defense in depth. If we handle flammable liquids or gases and an inherently safer design is not possible, we use some or all of the following:

• • • • • •

Prevent leaks by good design, construction, maintenance, and operation. Install automatic detectors so that leaks are detected promptly and people not required to deal with the leak can leave the area. Install remotely operated emergency isolation valves in places where leaks are most likely to occur or where a large quantity could leak. Remove all known sources of ignition. Minimize damage by installing fire protection. Passive equipment such as fire insulation is usually better than active equipment such as water spray turned on by automatic equipment. This is better than active equipment turned on by people. Provide fire-fighting equipment.

It is essential to carry out regular audits—tests and inspections to make sure that automatic equipment is in working order and those procedures have not lapsed.

Human Factors Engineers are interested in equipment, its failures, and ways of preventing them and often less interested in people. However, all systems involve both equipment and people. Engineers, whether they are designers, supervisors, or managers, therefore, should understand the way people react with equipment and why they sometime fail to act in the way we instruct them or expect them to act.

•

•

Some errors, usually called mistakes, occur because people do not know what to do. The intention was wrong. Employers should provide adequate training and instructions and should not write the sort of instructions that are designed to protect the writer rather than help the reader. However, for many instructions we write, problems will arise that are not covered by them and so people, particularly operators, should be trained in flexibility (i.e., the ability to diagnose and handle unforeseen situations). If instructions are not being followed, are they too complex? Can the job be simplified? Some errors, usually called violations or noncompliances, occur because someone knows what to do but makes a deliberate decision not to do it. Some violations occur because all people carrying out routine tasks tend to cut corners after a while. Many more occur because people think they know a better way

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Process Safety and Risk Management of doing the job. Note that if the instructions are wrong, noncompliance may achieve the intention. There is a fine line between showing initiative and breaking the rules. To prevent or reduce violations, we should do the following

• • • •

•

Explain the reasons for the instructions. We do not live in a society in which people will simply do as they are told. They want to know the reason why. If possible, simplify the job. If the correct method is difficult, an incorrect method will be used. Carry out checks from time to time to see that instructions are being followed and do not turn a blind eye if they are not. Some errors (mismatches) occur because the job is beyond the physical or mental ability of the person asked to do it, sometimes beyond anyone’s ability. For example, errors occur if people are overloaded, or underloaded, or asked to break well-established habits. We should change the plant design or method of working. The fourth category is the commonest—a momentary slip or lapse of attention. People know what to do, intend to do it, and are able to do it, but it slips their mind. Compared with mistakes, the intention is correct but is not fulfilled. They happen to everyone from time to time and cannot be prevented by telling people to be more careful or by telling them to keep their minds on the job. All we can do is to change the plant design or method of working so as to remove opportunities for error (or minimize the consequences or provide opportunities for recovery). We should, whenever possible, design inherently safer plants which can withstand errors (and equipment failures) without serious effects on safety (and output and efficiency).

Managers and designers as well as operators make errors, but because they usually have time to check their work, slip and lapses of attention are infrequent. Most of their errors are mistakes or violations.

Management Systems Some management systems have been discussed in earlier sections on risk assessment, hazard identification, and accident investigation. The following are also important: The preparation of equipment for maintenance: Many accidents have occurred because equipment was not isolated correctly, was not freed from hazardous materials, or was not correctly identified and the wrong equipment was opened up. Sometimes, procedures were poor; sometimes, they were not followed. The management of change: Many accidents have occurred because a change to plant, process, or organization had unforeseen effects. Before any change is made, it should be examined by professionally qualified people using HAZOP (or a simpler technique if the change is minor) and then inspected after completion to make sure that the intention has been followed and that the modification


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looks right. What does not look right is often wrong and should always be checked. Testing and inspection of equipment: All protective equipment is liable to fail and should be tested or inspected at regular intervals. When active equipment such as relief valves and interlocks fails, the failure is usually hidden and regular testing is necessary. If passive equipment such as fire insulation is missing, this is visible, but, nevertheless, it should be checked regularly. If 10% of the fire insulation on a vessel is missing, the rest is useless. The following equipment is often overlooked but should be tested or inspected regularly: • Drain holes in relief valve tailpipes. If they choke, rainwater will accumulate in the tailpipe. • Drain valves in tank bunds. If they are left open, the bund is useless. • Emergency equipment such as diesel-driven firewater pumps and generators. • Earth connections, especially the moveable ones used for earthing road tankers. • Fire and smoke detectors and fire-fighting equipment. • Flame arrestors. • Hired equipment. Who will test it, the owner or the hirer? • Labels are a sort of protective equipment. They vanish with remarkable speed and regular checks should be made to make sure that they are still there. • Mechanical protective equipment such as overspeed trips. • Nitrogen blanketing (on tanks, stacks and centrifuges). • Nonreturn valves and other backflow prevention devices, if their failure can affect the safety of the plant. • Open vents. These are the simplest possible sort of relief device and should be treated as relief valves. • Spare pumps, especially those fitted with auto-starts. • Steam traps. • Trace heating (steam or electrical). • Valves, remotely operated and hand-operated, which have to be used in an emergency. • Ventilation equipment. • Water sprays and steam curtains. All protective equipment should be designed so that it can be tested or inspected. Test results should be displayed for all to see, for example, on a board in the control room. Operators sometimes regard tests and inspections as a nuisance, interfering with the smooth operation of the plant. Training should emphasize that protective equipment is there for their protection and they should ‘‘own’’ it. Remembering the past: A most important system, discussed in the subsection Remembering the Message, is one to ensure that the lessons learned from past accidents, in our own and other companies, is not forgotten and that the information can readily be retrieved.

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Process Safety and Risk Management Introducing and maintaining systems: When systems are introduced or changed, they should be discussed with those who will have to operate them and not just sent to them through the mail. Discussions should start with descriptions of incidents that would not have occurred if the systems had been in operation at the time. These have much more impact than mere procedures and bring out the need for the changes. Discussions will allow the manager to check that the message has been received and understood and he may discover that it is impracticable or difficult to use in its present form. All systems are subject to a form of corrosion more rapid than that which affects the steelwork and can vanish without trace once managers lose interest. Continuous monitoring is necessary to make sure that systems continue in use. Limitations of systems: Some managers seem to believe that good safety management systems will ensure a safe plant. All the systems can do, however, is ensure that people’s knowledge and experience are applied systematically. If the staff lack knowledge and experience, then the systems are empty shells. People will go through the motions, but the output will be poor. Without a system, people will not achieve their full potential. Without knowledge and experience, systems will achieve nothing. This is a particular danger at times when companies are reducing manpower and experienced people are leaving. Senior managers should systematically assess the levels of knowledge and experience needed and ensure that they are maintained.

Audits We need audits of equipment and procedures by outsiders because of the following:

• • •

Those who work in a plant do not notice the hazards they see everyday. Auditors may have specialized knowledge and thus see hazards not apparent to others. Auditors have more time for investigation in depth than those who work regularly on a plant.

Safety auditing should not be a police activity; it is intended to help the local management, who may miss hazards through familiarity, ignorance, or lack of time. Auditors should pay particular attention to the following:

• • • •

The quality of the training and instructions and the knowledge and experience of employees. The procedures for preparing equipment for maintenance, controlling modifications, and testing protective equipment and whether or not these procedures are actually followed. Procedures for investigating accidents, passing on the lessons learned, and ensuring that they are not forgotten. Process hazards as well as mechanical ones.


• •

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Places where others do not look, behind, and underneath equipment. Although it may be a separate exercise, process hazards should be reassessed every few years in the light of new knowledge and new techniques.

Auditors (and managers) should visit the plant at night and at weekends, not just during the day.

The Measurement of Safety Whenever possible, we should provide a numerical measure of the success of each management function. Accident rates in good companies are now so low that the usual measure of safety, the lost-time accident rate, merely measures luck and the willingness of injured people to remain at work. In any case, it never measured process safety. Possible additional or alternative measures are as follows:

• • • • •

An index based on audit results. Unlike many other measures of safety, this one tries to detect falling standards before an accident occurs. A monthly summary of the cost of incidents. An annual report of the progress made in reducing inventories of hazardous substances. The number of faulty permits-to-work found by routine inspection. The number of faulty protective systems found by routine testing.

Mitigation Mitigation is the cornerstone of emergency management. It is the ongoing effort to lessen the impact disasters have on people and property. Mitigation involves keeping homes and populated areas away from industry, engineering process plants to be inherently safer, and creating and enforcing effective engineering codes to protect employees, the public, and the environment from potential process plant upsets and incidents. Mitigation is defined as ‘‘sustained action that reduces or eliminates long-term risk to people and property from hazards and their family and belongings are better protected from floods, earthquakes, hurricanes, and other natural hazards. They can be utilized to help business and industry avoid damages to their facilities and remain operational in the face of catastrophe. Mitigation technologies can be used to strengthen hospitals, fire stations, and other critical service facilities so that they can remain operational or reopen more quickly after an event. In addition, mitigation measures can help reduce disaster losses and suffering so that there is less demand for money and resources in the aftermath. In practice, mitigation can take many forms. It can involve actions such as the following:

• •

Promoting sound land use planning based on known hazards Buying flood insurance to protect your belongings

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• • • • • •

Relocating or elevating structures out of the floodplains Having hurricane straps installed to more securely attach a structure’s roof to its walls and foundation Developing, adopting, and enforcing effective engineering codes and standards Engineering process plants to be inherently safer Using fire-retardant materials in new construction Developing and implementing a plan in your business or community to reduce your susceptibility to hazards

In the multiple-barriers concept and development of inherently safer layers of protection, mitigation is a much lower-level activity and should be looked at only when all other measures in the inherently safer hierarchy are exhausted. For example, mitigation should be addressed only after the following options have been exhausted:

• • •

Inherent safety: These include inventory reduction (i.e., less chemicals stored or less in process vessels), substitution of a less hazardous chemical for one more hazardous, and use of lower temperatures and pressures. Engineering design: Examples are use of better seals or materials of construction, ensuring proper operating conditions and material purity, and installing dikes and spill vessels. Management: Examples include consistent operating policies and procedures, training for vapor release prevention and control, audits and inspections, equipment testing, maintenance program, management of modification and changes to prevent new hazards, and general plant security.

Some of the common mitigation techniques employed by process plants are as follows:

•

•

Early vapor detection and warning: Detection by sensors or personnel. Depending on the nature and extent of the chemical hazards, some plants may choose to employ very sophisticated sensor systems. For example, a pipeline company handling sour gas mixtures with very high H2S content decided to install an early warning H2S-sensing system. The system known, in the industry as Teledyne Geotech’s ‘‘LASP’’ [27,28] (Leak Alarm System for Pollutants) consisted of a semipermeable tubing which is laid above the sour gas pipeline under the ground. The tubing is capable of drawing air through it, which is analyzed for H2S contamination at regular intervals. Another common technique utilized to protect high-hazard pipelines is the installation of labeled warning ribbons approximately 1 ft below grade over the pipeline. Use of engineered and management systems to impede the progress of the released chemicals. Some of the engineered systems, which have been very effective in mitigation, include water sprays, water curtains, steam curtains, and air curtains. Management systems may include standing procedures to deliberately ignite explosive clouds, procedures for forced dilution of contaminant, and procedures to mitigate or suppress released chemicals by the use of foams and other suppressants.


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Isolation by distance of a chemical process from on-site and off-site surrounding populations is generally a very effective consequence–mitigation measure. Separating the process from vulnerable populations affords both attenuation of the effects and time to provide emergency response. The isolation distances needed to appreciably reduce blast effects and impact from toxic releases is significantly large, whereas this type of mitigation measure is not useful for the protection of on-site personnel. Other hazard mitigation measures should be used for protection of on-site personnel. For example, explosion hazards may warrant the construction of blast-resistant buildings or blast walls. Toxic release hazards may require the availability of shelter-in-place facilities or escape respirators. Some process plants use consequence modeling in deciding the layout of the plant. For example, a hazards analysis early in the design stage may identify one particular unit as having the greatest potential for a toxic release and that unit may then be located on the site as far as possible from off-site neighbors, perhaps considering prevailing directions as well. Process integrity may also be addressed in the engineering design. Process integrity involves the chemistry of plant design and operation. Mitigation after loss of containment can also be effective and usually must be provided for in the process design stage. Secondary containment by double-walled piping or double-walled vessels may be needed. Dikes, curbs, and trenches leading away from storage vessels to strategically located impoundments can be used to reduce the rate of evaporation, help keep the liquid source of the vapor away from the most sensitive areas of the plant, and limit the extent of emergency response activities. Mitigation measures such as active and passive scrubbers, stacks, flares, catch tanks for vapor–liquid separation, incinerators, absorbers, adsorbers, and condensers are used widely for reducing the impact after loss of containment. CCPS’s Guidelines for Vapor Release Mitigation [29] provides detailed discussions on these mitigation techniques.

Response Emergency response plans for process plants should be developed in accordance with applicable governmental regulations and operating company requirements. Written emergency response procedures, accident investigation protocol and procedures, and repair procedures should be prepared, and the appropriate operating personnel should be trained in their proper use. Results or risk assessment studies and hazard zone calculations should be used in the formulation of emergency response actions contained in the emergency plans. In addition to being quickly and effectively warned of a dangerous situation, personnel need to know ahead of time the best response to minimize their chances of being affected by the vapors. Questions that should be settled in advance include emergency shutdown criteria, incident commander designation, and the roles of incident commander and other emergency personnel. Other issues that should be settled in advance include circumstances which would warrant shelter-in-place for the employees and the affected public.

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Process Safety and Risk Management Communicating warnings to potentially exposed personnel is essential in an emergency. Warnings can be communicated through the use of public address systems, alarm systems, and sirens. Recently, many companies are also moving toward the installation of automatic telephone dialing and alerting systems or communicating hazard warnings to downwind personnel. Preprogrammed computers can be used to dial thousands of preselected numbers in very short periods of time. Every plant should establish clearly defined procedures for emergency shutdown of equipment. The procedures should clearly indicate what constitutes an emergency situation. The extent and nature of steps taken to bring the process back under control should also be clearly indicated. Finally, the conditions which would unambiguously require a shutdown should be spelled out. Personnel authorized to make the decision to shutdown should be aware of their responsibilities and the role of other personnel.

Technology Advances Advances in the understanding of chemical hazards have led to the development of new technology in the arena of process design, equipment, and risk management. A discussion of three major areas of development is given here.

Relief Valve Sizing and Overpressure Alternatives Recent research on relief valve sizing and overpressure protection alternatives has focused on the development of validated engineering design procedures for the proper sizing of safety relief valves for systems, which involve two-phase flows of viscous fluids. Systems which are being considered include single-phase viscous liquids and gas flows, ‘‘frozen’’ (e.g., air–liquid) two-phase flows of gases and viscous liquids, and flashing flows of viscous and nonviscous liquids.

Reactive Chemistry In the Reactive Chemistry arena, calorimeters are being used increasingly for studying the thermal behavior of reactive systems. One such calorimeter is the Reactive Systems Screening Tool (RSST), which is designed for rapid measurement of thermal behavior of small samples (10 cm3) for temperatures up to 400°C and pressures to 500 psia. Another apparatus is the Automatic Pressure Tracking Adiabatic Calorimeter (APTAC) for detailed analyses of thermal behavior of larger samples (up to ⬃130 cm3) for temperatures up to 450°C and pressures up to 2000 psia. In this calorimeter, closed-cell sample pressures are continuously matched by an external pressure of nitrogen so that sample cells of low mass and therefore low thermal inertia can be used for highly sensitive measurements of sample thermal behavior. Other advanced features of the APTAC include in situ additions to the sample cell of reactants or catalysts with a high-pressure syringe pump.


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During typical experiments with each calorimeter, the sample temperature is measured during temperature scans when the temperature is increased at a steady rate or held isothermally (in the APTAC). Thermal energy released or absorbed by the sample (as defined by sample temperature changes of ⱖ0.04°C/min) is measured by the calorimeter as a nearly adiabatic excursion from the thermal scan baseline. For each sample, the thermal peaks can be identified and measured from ambient up to 450°C (up to 400°C in the RSST). With this experimental capability, investigations of thermal behavior of wide ranges of reactive systems and systems of questionable chemical compatibility can be performed, which, in turn, is used to design safe processes and choose safe operating conditions.

Safety Integrity Levels Industry is moving toward the use of high-integrity protection systems to reduce flare loading and alleviate the need to upgrade existing flare systems when expanding facilities. In the process industry, a key safety consideration is the control and response to overpressure situations. Industry standards from the American Petroleum Institute (API) and American Society of Mechanical Engineers (ASME) provide criteria for the design of vessels and the protection of these vessels from overpressure. Traditionally, pressure relief valves and flares were used to handle the relieving of vessels in the worst credible scenario. Flare loading calculations gave no credit for operator intervention, fail-safe equipment operation, or trip systems. In many communities and countries around the world, the belt is tightening on the venting and combustion of gases. It is simply not acceptable to flare large volumes of gas. In addition, the cost of designing and installing large flare systems has continued to rise. API 521 and Case 2211 of ASME Section VIII, Division 1 and 2, provide alternatives in the design of overpressure protection systems. These alternatives revolve around the use of an instrumented system that exceeds the protection provided by a pressure relief valve and flare system. ASME Code Case 2211, approved in 1996, sets the conditions under which overpressure protection may be provided by an instrumented system instead of a pressure relief valve (PRV). The ruling is intended to enhance the overall safety and environmental performance of a facility by utilizing the most appropriate engineered option for pressure protection. Although there are no specific performance criteria in the Case Code, the substitution of the high-integrity protection systems for the pressure relief valve should provide a safer installation. Consequently, the substitution is generally intended for limited services where the PRV may not work properly due to process condition (e.g., plugging, multiple phases, etc.). The overpressure protection can be provided by a safety instrumented system in lieu of a pressure-relieving device under the following conditions: One of the most important criteria for safety instrumented system (SIS) design is the requirement that the User assign and verify the safety integrity level (SIL) for the SIS [30]. The assignment of SIL is a corporate decision based on risk management philosophy and risk tolerance. The SIS should be designed to meet a safety integrity level, which is appropriate for the degree of hazard associated with the process upset. Safety integrity levels per draft IEC 61508 [31,32] and ANSI/ISA S84.01 [33–37] are designated in Table 7.

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Process Safety and Risk Management TABLE 7 Safety Integrity Levels Safety integrity level IEC 61508: 4 ISA S84: 3 2 1

Availability required

Probability to fail on demand

1/PFD

⬎99.99%

10⫺5 to 10⫺4

100,000–10,000

99.90–99.99% 99.00–99.90% 90.00–99.00%

10⫺4 to 10⫺3 10⫺3 to 10⫺2 10⫺2 to 10⫺1

10,000–1,000 1,000–100 100–10

Industrial Hygiene and Toxicology As the world becomes more industrialized in an attempt to increase the quality of life, more and more environmental problems will require the attention of engineers, managers, and planners. In the coming years, findings and advances on chemical toxicity will require the implementation of stringent industrial hygiene standards. Although there are many toxic effects, both acute and long term associated with chemicals, one of the most dreaded is cancer. It is quite apparent that industrial hygiene requirements in the coming years will be dictated to a large extent by the upcoming findings on carcinogenesis and mutagenesis. By one estimate, there are 500 new chemicals marketed each year [38]. Thus, determining the toxicological effects of these chemicals and, as a result, developing industrial hygiene programs to protect people from these effects will command significant attention and resources from process safety personnel. Although workers are often exposed to contaminant mixtures, exposure regulations do not take into account the effects of the various types of contaminant interactions capable of modifying toxicity. To remedy this situation, the scientific research and development of databases on the toxicity of mixtures are needed. With this information, health and safety specialists will be able to quantify contaminant interactions for any given situation. Although substantial progress has been made in the United States toward improving worker protections since 1970 (largely a result of occupational safety and health research), workplace hazards continue to inflict a tremendous toll in terms of human and economic costs. Clearly, there is much work to be done. The practical impact of the toxicological and industrial hygiene research programs on the workplace largely depends on the actions of employers, employees, and partners in governmental agencies, industry, labor, academia, and community organizations. The stated objectives in this area include target levels of improvements in work-related conditions. Examples are reducing work-related deaths and injuries, reducing lost work days and incidences of cumulative trauma and skin disorders, and increasing the number of workplaces with rehabilitation and safety and health programs. The work of governmental agencies and research organizations has had and will continue to have an impact on improving health and safety at the workplace; therefore, it will help address many of the issues related to workrelated hazards, injuries, illnesses, and deaths (such as musculoskeletal problems,


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skin diseases, violence in the workplace, employee stress, and back injuries) as well as categories of workers and prevention strategies for mine workers, farm workers, and adolescents. In addition, surveillance efforts will assist the development of comprehensive databases, thereby helping to establish baseline and trend information in the occupational safety and health area.

Future Developments Increasingly, the process safety requirements for chemical plants will become more and more stringent. In addition, the pressure to operate safely from the point of view of competitiveness and profitability will also keep increasing. Finally, the public outcry for improved safety performance also creates significant pressure on the industry. In fact, in future processes, safety performance will quite likely be dictated by national goal setting. This would require the establishment of a baseline assessment of the status of process safety incidents. Given a baseline assessment, National Chemical Safety Goals can be established, with the identification of activities necessary to accomplish the goals and the development of a measurement system to measure progress toward the goals. Regulatory programs and industrial standards and practices in the United States have quite often been reactive (i.e., in response to catastrophic accidents or other events). The pros and cons of establishing national process safety goals and evaluation approaches include the following: 1. Stakeholder consensus on national chemical (process) safety goals 2. Identification of where we want to be and by when in relation to national chemical safety goals 3. List of activities that need to be implemented to accomplish Step 2 above 4. Agreement on some common metrics for measurement of progress toward national chemical safety goals

Summary and Conclusions The industrial revolution brought prosperity and, along with it, the use of hazardous processes and complex technologies. Growing economies and global competition has led to more complex processes involving the use of hazardous chemicals, exotic chemistry, and extreme operating conditions. As a result, a fundamental understanding of the hazards and associated risks is essential. Process safety and risk management requires the application of the basic sciences and a systematic approach. Recent advances, such as overpressure protection alternatives and reactive chemistry, allow safer design and operation of processes. In the multiple-barriers concept, plants are designed with several layers, so that an accident would require the failure of several systems. Another novel approach to

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Process Safety and Risk Management process safety and risk management is to consider various actions in a descending hierarchical order. Inherently safer design consideration should be first in the hierarchy, followed by prevention systems, mitigation, and response. The success of these systems is dependent on the fundamental understanding of the process and the associated hazards. Chronic as well as catastrophic consequences resulting from toxic and flammable substances can be reduced and/or eliminated through appropriate design and operating practices. In the end, progress toward the improvement in safety performance can be measured only by a reduction in occupational injuries, illnesses, and fatalities. In fact, measurable progress has been made in the period 1970 to 1995, during which the rate of workplace fatalities fell by 78% and the number of workplace deaths has declined by 62%. We have also seen a 25% decline in the rate of occupational injuries and illnesses from 1973 through 1994. These reductions are the result of the combined efforts of all the partners in occupational safety and health: industry, labor, academic researchers, National Institute of Occupational Safety and Health, Occupational Safety and Health Administration, Mining Safety and Health Administration, state and local agencies, and others. No single partners can claim exclusive credit for the progress. Thus, if further progress is to be made, all of the partners must act—from identifying the causes of disease and injury through controlling or eliminating the hazards or exposures at the worksite.

References

1. M. Connors, ‘‘The Battle for Industrial Safety,’’ Fortune, 116[C-P] (August 4, 1997). 2. D. A. Crowl and J. F. Louvar, Chemical Process Safety: Fundamentals with Applications, Prentice-Hall, Englewood Cliffs, NJ, 1990. 3. Center for Chemical Process Safety, Guidelines for Safe Automation of Chemical Processes, American Institute of Chemical Engineers, New York, 1993. 4. American Petroleum Institute, Recommended Practice 750: Management of Process Hazards, API, Washington, DC, 1990. 5. ‘‘Notice of Proposed Rule Making on Process Safety Management of Highly Hazardous Chemicals’’: 29 CFR 1910.119, Federal Register, Washington, DC, July 17, 1990. 6. Chemical Manufacturers Association, Resource Guide for Implementing the Process Safety Management Code of Practices, Chemical Manufacturers Association, Washington, DC, 1990. 7. ‘‘Final Rule on Process Safety Management of Highly Hazardous Chemicals’’: 29 CFR 1910.119, Federal Register, Washington, DC, February 24, 1992. 8. Norwegian Petroleum Directorate, ‘‘Safety Evaluation of Platform Conceptual Design,’’ Stavanger, Norway, 1981. 9. European Community Directive, ‘‘On the Major Accident Hazards of Certain Industrial Activities,’’ 82/501/E, J. Eur. Community, L230 (June 1982). 10. Offshore Installation (Safety Case) Regulation 1992, Health and Safety Executive, London, UK, 1992. 11. ‘‘Techniques for Assessing Industrial Hazards,’’ World Bank Technical Paper #55, Washington, DC, 1988. 12. Major Hazard Control, a Practical Manual, International Labour Office, Geneva, 1988.


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13. Center for Chemical Process Safety, Guidelines for Chemical Process Quantitative Risk Analysis, American Institute of Chemical Engineers, New York, 1989. 14. R. J. Lewis (ed.), Sax’s Dangerous Properties of Industrial Materials, 9th ed., John Wiley & Sons, New York, 1996. 15. P. G. Urben (ed.), Bretherick’s Handbook of Reactive Chemical Hazards, 5th ed., Butterworth-Heinemann Boston, 1995. 16. Dow Chemical Company, Dow’s Fire and Explosion Index Hazard Classification Guide, 7th ed., American Institute of Chemical Engineers, New York, 1994. 17. Dow Chemical Company, Dow’s Chemical Exposure Index Guide, American Institute of Chemical Engineers, New York, 1994. 18. Center for Chemical Process Safety (CCPS), Guidelines for Hazard Evaluation Procedures, 2nd ed., with Worked Examples, American Institute of Chemical Engineers, New York, 1992. 19. G. Wells, Hazard Identification and Risk Assessment, Institution of Chemical Engineers, Rugby, Warwickshire, UK, 1996. 20. Center for Chemical Process Safety (CCPS), Guidelines for Use of Vapor Cloud Dispersion Models, 2nd ed., American Institute of Chemical Engineers, New York, 1996. 21. Center for Chemical Process Safety (CCPS), Guidelines for Evaluating the Characteristics of Vapor Cloud Explosions, Flash Fires, and BLEVES, American Institute of Chemical Engineers, New York, 1994. 22. Center for Chemical Process Safety (CCPS), Guidelines for Consequence Analysis of Chemical Releases, American Institute of Chemical Engineers, New York, 1999. 23. G. E. DeVaull, J. A. King, R. J. Lantzy, and D. J. Fontaine, Understanding Atmospheric Dispersion of Accidental Releases, American Institute of Chemical Engineers, New York, 1995. 24. The Netherlands Organization for Applied Scientific Research (TNO), Methods for the Calculation of Physical Effects, Part 1 and 2 CPR-14, 3rd ed., SdU Uitgevers, The Hague, 1997. 25. F. P. Lees, Loss Prevention in the Process Industries, 2nd ed., Butterworth-Heinemann, Boston, 1996. 26. H. G. Fisher, H. S. Forrest, S. S. Grossel, J. E. Huff, A. R. Muller, J. A. Noronha, D. A. Shaw, and B. J. Tilley, Emergency Relief System Design Using DIERS Technology, American Institute of Chemical Engineers, New York, 1992. 27. M. Mannan, D. B. Pfenning, and C. D. Zinn, ‘‘Sour Gas Pipeline—1: Risk-Analysis Procedures Ensure System Safety,’’ Oil Gas J. 83–87 (June 3, 1991). 28. M. Mannan, D. B. Pfenning, and C. D. Zinn, ‘‘Sour Gas Pipeline—Conclusion: Line, Weather Conditions Among Variables to Determine Public Risk,’’ Oil Gas J., 34– 35 (June 10, 1991). 29. Center for Chemical Process Safety (CCPS), Guidelines for Vapor Release Mitigation, American Institute of Chemical Engineers, New York, 1988. 30. A. E. Summers, ‘‘Techniques for Assigning a Target Safety Integrity Level,’’ ISA Trans., 37, 95–104 (1998). 31. IEC 61508, 65A/255/CDV, ‘‘Functional Safety of Electrical/Electronic/Programmable Electronic Safety Related Systems, Parts 1, 3, 4, and 5,’’ International Electrotechnical Commission, Final Standard, December 1998. 32. IEC 61508, 65A/255/CDV, ‘‘Functional Safety of Electrical/Electronic/Programmable Electronic Safety Related Systems, Parts 2, 6, and 7,’’ International Electrotechnical Commission, Final Draft International Standard, January 1999. 33. ‘‘Safety Instrumented Systems (SIS)—Safety Integrity Level (SIL) Evaluation Techniques, Part 1: Introduction,’’ TR84.0.02, Draft, Version 4, March 1998. 34. ‘‘Safety Instrumented Systems (SIS)—Safety Integrity Level (SIL) Evaluation Tech-

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35.

36.

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niques, Part 2: Determining the SIL of a SIS via Simplified Equations,’’ TR84.0.02, Draft, Version 4, March 1998. ‘‘Safety Instrumented Systems (SIS)—Safety Integrity Level (SIL) Evaluation Techniques, Part 3: Determining the SIL of a SIS via Fault Tree Analysis,’’ TR84.0.02, Draft, Version 3, March 1998. ‘‘Safety Instrumented Systems (SIS)—Safety Integrity Level (SIL) Evaluation Techniques, Part 4: Determining the SIL of a SIS via Markov Analysis,’’ TR84.0.02, Draft, Version 4, March 1998. ‘‘Safety Instrumented Systems (SIS)—Safety Integrity Level (SIL) Evaluation Techniques, Part 5: Determining the PFD of SIS Logic Solvers via Markov Analysis,’’ TR84.0.02, Draft, Version 4, April 1998. T. F. Yen, Environmental Chemistry: Essentials of Chemistry for Engineering Practice, Prentice-Hall, Englewood Cliffs, NJ, 1999. M. SAM MANNAN DENNIS HENDERSHOT TREVOR A. KLETZ

Introduction to the Selective Catalytic Reduction Technology

Introduction Overview The selective catalytic reduction (SCR) process has been originally developed for reducing oxides of nitrogen (NOx). The process is being discussed within the framework of air-pollution control policies and practices of industrialized countries, accounting for about 20% of the world population and generating 80% of the global combustion air-pollution. The United States with approximately 5% of the world population accounts for about 28% of the world’s fossil fuel consumption (International Energy Agency, Paris, France), more than twice the per-capita consumption of other major industrial countries. Stringent Energy Conservation and Emission Reductions Policies have a much longer tradition in Europe and Japan than in the United States. In the United States, due to low-cost energy, excess fuel is often used to reduce NOx of IC engines, which accounts for over 50% of the United States’ total NOx inventory, by a 4-degree time-retard timing. This at the expense of major increases in CO emissions, fuel consumption, and carcinogenic particulate matter (PM) emission. (Source: BACT best available control technology of U.S. EPA, Cal.-ARB, and SCAQMD). The SCR process has been designed to reduce oxides of nitrogen (NOx), a precursor of regulated low-level ozone (O3) emission. The ozone is formed primar-

Selective Catalytic Reduction

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ily by the two precursor NOx and volatile organic compounds (VOC) in the presence of sunlight through photosynthesis. However, advanced SCR developments have been achieving VOC/HC (hydrocarbons) and particulate matter (PM) emission reductions simultaneously with the NOx reduction. Also, in recent years, European application engineering technology developments have broadened the application of SCR technology. Originally developed in the 1970s for stationary utility power generation equipment, the SCR process can be used to reduce emissions of practically all types of combustion equipment, including mobile applications such as heavy-duty diesel (HDD) trucks, portable generation sets, diesel locomotives, coastal and ocean-going vessels, earth-moving equipment, and so forth. Today, SCR technology is being used in retrofit and integrated emission control systems. The SCR combustion pollution control technology is able to overcome the target conflict of IC combustion equipment, increasing the NOx emission by reducing the fuel consumption or increasing the PM and CO emission by increasing fuel consumption through fuel-based emission controls (see Fig. 1). The earliest SCR applications were located primarily in oil-burning utility boiler facilities in Japan in the 1970s. Then, in the 1980s, European SCR applications for hard-coal-burning utility and oil- and gas-burning industrial boilers, and stationary gas, diesel, and dual-fuel IC engines were added. In the United States, gas turbine applications were introduced in the 1980s and many coal-burning utility power plants require SCR retrofit system installations by 2004. Figure 2 shows high-dust and low-dust utility boiler applications. However, because the number of utility plants are relatively few in comparison to the number of other nonutility SCR applications to be considered for OEM and retrofit applications in future, accounting for over half of the total U.S. NOx pollution inventory, the nonutility and the new mobile combustion sources are being emphasized in this article. The SCR utility applications have to compete with pollution control processes,

FIG. 1 The target conflict of IC equipment emission reduction versus fuel consumption. (Source: German Engineering Society (VDI), May 1993.)

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FIG. 2 Utility boiler applications. (Courtesy of Siemens.)

such as the selective noncatalytic reduction process (SNCR), featuring low capital investment but high operating costs. This process is limited to a small operating temperature window of usually 1350–1550 F (750–900°C) and requires an ammonia or urea consumption of up to four times the stoichiometric requirements of the SCR process. In the United States, natural-gas- and fuel-oil-burning industrial boilers and IC engines and turbines were equipped with SCR systems in a few regional nonattainment areas for ozone since the 1980s and gas- and coal-burning utility boilers since the 1990s. In Europe, due to the development of a diesel SCR catalyst with an operating temperature window of 300–1020 F (150–550° C), air quality regulators cooperated with the trucking industry and the SCR equipment industry in the development and field testing of SCR systems for heavy-duty diesel truck engine applications in the 1990s. These truck field tests have lasted for some 4 years, accumulating approximately 3 million road and highway miles with 20 Class 8 type HDD trucks. The SCR systems will go to market in 2002. Today, the range of SCR applications is rather broad, as shown by Fig. 3: From (clockwise) a coal-burning utility plant to HDD truck, portable generator set, vessel, railroad, construction equipment, pipeline pump station, and gas turbine applications. Today, the SCR technology allows the simultaneous reduction of combustion emission of 70–95% NOx, 20–50% PM, and 85–95% VOC/HC at minimum fuel consumption and CO2 emission. It may be interesting to note that the emission reduction of VOC include air toxins such as aldehydes and polycyclic aromatic hydrocarbons (PAH), which are almost completely oxidized in the SCR process. Hence, SCR technology is considered the most promising, cost-effective technology available to reduce combustion emissions, enabling air quality regulators worldwide to achieve more drastic emission reductions of combustion equipment than thought possible just a few years ago. Various cost-effectiveness calculations showed annualized costs of ⬍$500 to ⬃1500 per annual ton of NOx reduction, depending on the type and size of the SCR equipment. The capital investment cost of SCR systems are ⬃US$ 10 to 50 per BHP h, depending on custom engineering requirements and the number of systems of the same design fabricated (Siemens).


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FIG. 3 The Range of SCR applications. (Courtesy of Siemens.)

Environmental Policies After the United Nations summits on the environment in the 1990s in Rio, Brazil and Kyoto, Japan, environmental policy-makers around the world have been focusing on the reduction of emission generated by combusting fossil fuel.

•

•

Health effects of NOx, PM, and VOC/HC have been the driving force for air quality legislation in the past. The air quality regulations, however, were developed in a piecemeal approach: first NOx, then VOC/HC to reduce ground-level ozone, and, finally, PM. This has been a major disadvantage for SCR technology applications, as no considerations were given to the capability of reducing VOC and PM simultaneously with NOx at no extra cost. Global warming is becoming an increasing concern domestically and internationally, considering changes in weather pattern, shorelines of the ocean, and so forth. The industrialized countries must reduce their fossil-fuel combustion emission substantially to allow developing nations to generate more power for their growing economies. The various emission inventory data for the United States for NOx indicate that the leading sources of combustion emission are from on-road and off/nonroad emission sources. Gasoline engines, at the expense of higher fuel consumption and N2O (laughing gas) secondary emission, have been equipped with three-way catalyst systems for many years. For the most fuelefficient device in converting fossil fuel into useful energy and power, the diesel engine, no cost-effective, commercially available solution to reduce NOx, PM, and VOC/HC emissions substantially and simultaneously existed until the early 1990s. Therefore, as mentioned above, fuel is often used today to reduce emissions, increasing the fuel consumption of engines and other combustion equipment and, with it, global warming gas emissions. Increases of 10–13% in CO2

98


FIG. 4 Target conflict: reduction of NOx versus increase in fuel consumption, CO2, and PM. (Courtesy of Siemens.)

•

and up to 70% of PM emission of diesel engines were reported at public Cal. ARB meetings in 1999 and 2000. Combustion emission reductions through fuel savings and internal combustion equipment design modifications such as electronic engine management systems, unit electronic–hydraulic injection, intercooling, and exhaust gas recycling is limited. In Fig. 4, the target conflict of the U.S. EPA NOx emission reductions versus exponential increases in fuel consumption, CO, VOC/HC, and PM emission of Class 8 HDD truck engines, is shown by reducing NOx emissions of engines by modifications such as U.S. EPA’s BACT mentioned earlier. Therefore, to avoid higher fuel cost and to achieve future U.S. EPA emission reduction goals beyond 2002 and 2004, exhaust gas aftertreatment technologies will be required in mobile and other applications.

Environmental Regulations In 1970, the U.S. Ambient Air Quality Act was passed and then amended in 1990, allowing a maximum ground-level ozone concentration of 0.12 parts per million (ppm). The U.S. National Ambient Air Quality Standard (NAAQS), however, was never met throughout the United States. Serious nonattainment areas for ozone (O3) such as in the Northeast and in California still exist and new ones were added recently. In 1997, the U.S. EPA further reduced the ground-level ozone standard due to late health data and increased health concerns to 0.08 ppm ozone at an 8-h average and introduced a new PM 2.5 standard of 65 µm/m3 at a 24-h average. That increased the nonattainment areas in the United States substantially (Fig. 5). The various state implementation plans for combustion emissions in the United States, therefore, call for substantial further NOx reductions of the U.S. inventory of approximately 25 million tons of NOx per year (U.S. EPA).


99

FIG. 5 Nonattainment areas in the United States for ozone (dark areas) based on the EPA’s new standards for ground-level ozone and PM 2.5.

In addition to major cities in California and the Northeast, cities such as Atlanta, GA, Chicago, IL, Dallas, TX and Houston, TX became serious or extreme nonattainment areas. East coast states, after years of complaining about Mid-west combustion pollution transfers to the East, finally received an ozone transfer regulation for about 20 states, requiring midwest states to reduce NOx substantially. According to NESCAUM data presented at the ICAC Forum 1998, the NOx emission in the midwest was caused by utilities (⬃33%), by independent power producers (14%), and by mobile equipment (on-road 34% and off/nonroad 14%). Other estimates (i.e., California) are more like 30% each for stationary, on-road, and nonroad combustion equipment after 20 years of stationary emission source regulations in that State. Figure 6 shows the different emission source categories by area in Texas in 1996.

Technology As stated earlier, SCR technology has been advanced considerably in Europe in recent years; however, the basic SCR process is still the same. The SCR process equipment layout is pictured in Fig. 7, using an engine as an example for combustion equipment.

100


FIG. 6 NOx emission by source category in tons per year at selected areas of Texas in 1996. Example: Houston, clockwise: 2% area, 20% nonroad, mobile, 25% on-road, and 53% stationary emission sources. (Source: TNRCC, 2000.)

FIG. 7 The SCR process, using a diesel engine SCR application as an example. (Courtesy of Siemens.)


101

The exhaust gas generated by a diesel engine passes from the engine through the exhaust pipe or duct, through the SCR reactor and other exhaust gas system components such as the muffler into the atmosphere. The exhaust gas duct, leading to the reactor, houses the gas flow straightener, turning vanes, static mixers, and some sensor probes. The SCR reactor houses the SCR catalyst and other sensor and instrumentation probes. Depending on the exhaust gas temperature, mass flow, raw or uncontrolled, and permitted or controlled NOx concentration and other variables, a certain amount of a reducing agent is injected and homogenously mixed upstream of the catalyst bed. There are several reducing agents in use today. Anhydrous ammonia (NH3), a toxic, hazardous, and flammable gas, is used primarily in stationary industrial and utility applications. A safer ammonia reducing agent, aqueous ammonia, containing 25–29% ammonia gas in demineralized water, was introduced to SCR applications in the 1980s. The mobile SCR system application developments of the 1990s, however, required a much safer reducing agent, aqueous urea, from which ammonia is generated through hydrolysis, yielding two NH3 and one CO2 in the exhaust duct upstream of the SCR catalyst. When subsequently passing through the SCR catalyst bed, the reducing agent NH3 reacts with the NOx to yield molecular nitrogen (N2) and water vapor (H2O) with very limited ammonia slip, typically in the 2–30-ppm range as secondary emission. Some typical chemical reactions of the SCR process are shown by Fig. 8a. Figure 8b shows the reversible side reactions and their equilibrium at threshold temperatures (EESI/Steuler, a leading SCR system supplier) at a specific chemical exhaust gas composition.

Secondary Reactions and Emissions The effects of high SOx levels in the exhaust gas is given in Fig. 8b. A European university study on the effects of high SOx levels in the exhaust gas duct concluded that temperature and SO2 ⬎ SO3 control is key to avoid ammonia salt emissions. Figure 8b shows a few of the reversible side reactions and their equilibrium threshold temperatures (EESI/Steuler), using an exhaust gas of a specific chemical composition: 72% N2, 15% CO2, 10% H2O, 3% excess O2, 200 ppm NO, 10 ppm NO2, 1000 ppm SO2, 50 ppm SO3, and 200 ppm HCL. Ammonia slip has been a concern of health experts and air quality regulators when considering SCR exhaust emission reductions. United States air quality regulators for gas turbine SCR applications have imposed extremely costly low levels of NH3 slip emission limits of 2–5 ppm. However, pollution control experts point out that this ‘‘NH3-slip hype’’ is not based on facts: European and U.S. NH3 Emission Inventory Studies in the Netherlands, Germany, and California show that, depending on the life stock concentration, 60–75% of the total local NH3 emission is caused by animal droppings. Also, depending on weather conditions and the type and intensity of fertilizer usage, the area NH3 emission may be as high as 5– 30%. The industrial applications such as air conditioning, nitric acid production, and SCR-based NH3 slip emission do not account for more than 0.5–2% of the total NH3 emission inventory (Bavarian-EPA, Cal. ARB, and other studies). In a U.S. study, a well-supported estimate of a SCR-based NH3 slip rate of 2–5 ppm for 200,000 MW utility applications would just cause an NH3 emission increase

102

Selective Catalytic Reduction Reducing Agent Urea: Ammonia generation through hydrolysis in the exhaust or flue gas, upsteam of the SCR catalyst. 1. (NH2)2 CO ⫹ H2O

→ 2 NH3 ⫹ CO2

NOx Reduction with Ammonia 2. 3. 4. 5.

4NO ⫹ 4NH3 ⫹ O2 6NO ⫹ 4NH3 2NO2 ⫹ 4NH3 ⫹ O2 6NO2 ⫹ 4NH3

→ → → →

4N2 5N2 3N2 7N2

⫹ ⫹ ⫹ ⫹

6H2O 6H2O 6H2O 12H2O

(a) Effects of high SOx levels in the exhaust gas Temperature (°C) 1. 2. 3. 4. 5. (b)

22 219 209 113 88

Chemical Reactions 2NH3 ⫹ H2O ⫹ SO2 2NH3 ⫹ H2O ⫹ SO3 NH3 ⫹ H2O ⫹ SO3 NH3 ⫹ HCl NH3 ⫹ H2O ⫹ (1/2)O2 ⫹ 2NO2

i i i i i

(NH4)2SO3 (NH4)2SO4 NH4HSO4 NH4Cl 2NH4NO3

(c) FIG. 8 (a) Typical chemical reactions of the SCR process. (b) Equilibrium threshold temperatures at which ammonia salts is formed/dissolved. (c) NH3 /NOx ratio and the effects of overinjection and underinjection of ammonia. (Courtesy of L. Pruce.)

of less than 10,000 tons per year equal to less than 1% of the total U.S. NH3 emission inventory.

System Controls The component selection will vary depending on the type of combustion equipment, emission reduction, control/monitoring system requirements, and so forth.


103

Basically, two types of reducing agent injection control are being used. The first is the open-loop, feed-forward control with a computer-based ‘‘map,’’ a correlation function of NOx emission generated at certain combustion equipment loads, requiring a calculated amount of reducing agent to achieve the permitted NOx emission rate. Such a system is also called a Predictive Emission Monitoring System (PEMS). Maximum reduction rates at minimum unreacted ammonia (slip) can be achieved by adding to the rapidly responding open-loop system a slower closedloop system. The feedback systems are, in most cases, certified systems, incorporating continuous emission monitoring systems (CEMS) or sensors.

SCR System Design Considerations The following basic design considerations are common to all SCR systems:

• • •

Proper exhaust gas flow and temperature distribution at the front face of the catalyst bed Proper selection of the catalyst material formulation, the catalyst bed, and housing design Proper selection of materials, system components, and control system design

The SCR system component selection and system configuration may vary considerably depending on the application. Utility applications for large boilers or gas turbines are therefore custom engineered. Also, Independent Power Producer (IPP) and industrial applications such as cogeneration and standby generation sets are still primarily custom engineered. However standardized, volume-produced small gas turbine and engine SCR applications will be more cost-effective and thus more common in future. The new mobile SCR applications for diesel and lean-burn gas engines will be such standardized and volume-produced SCR systems. This industrial approach, developed and adopted for the first time in the HDD truck program by Siemens and the European trucking industry, is already benefiting the custom-engineered SCR systems as well.

SCR System Component Summary

The Exhaust or Flue Gas Piping or Duct This is usually heat insulated, made of heat-resistant nonscaling steel such as molybdenum alloys, and connects the combustion equipment with the SCR reactor. Such inlet ducts with the SCR reactors downstream are pictured in Fig. 9. The inlet duct or piping houses the gas flow straightener, flow turning vanes, sensor probes, and injection spray heads or nozzle as well as the static mixers. Only in very special cases, bypass valves are incorporated in ducts to cope with certain operating conditions such as cold start-ups, alternate fuels, and so forth, which could harm the SCR catalyst of earlier developments. An even flow rate

104


(a)

(b) FIG. 9 (a) SCR inlet duct and catalyst location, gas turbine application. (b) SCR inlet duct and catalyst location, stationary engine application.

and temperature distribution at the front face of the catalyst bed are most crucial and prerequisites of any well-performing SCR system design. As a case in point, in a gas turbine project in southern California, even major improvements in the ammonia/exhaust gas mixing system and a major increase in the catalyst volume could not solve the original problem. The uneven gas flow rate and temperature distribution at the front face of the catalyst still deviated by 20–40% due to improper turning vanes and lack of static mixers that the contractor had refused to install. To obtain a Permit to Operate, the permitted NOx reduction rate was finally negotiated and increased by the local regulator without hearing expert witnesses. Figure 10 pictures such a flow modeling case and how an even gas flow can be achieved through turning vanes and a static mixer in front of the SCR catalyst bed.

The Reactor Housing and Stack Duct These are normally heat insulated, made of heat resistant nonscaling molybdenum alloy steel rather than stainless steel, which is sensitive to stress corrosion in welded areas. They house the SCR catalyst and any oxidation catalyst or sound deadening material as may be required by the specific application. In addition, they house instrumentation probes for monitoring operating conditions and emission


105

FIG. 10 Flow modeling: before and after computer and scale modeling, improving the uneven (32%) to a more acceptable (9%) flow rate deviation at the front face of the catalyst bed.

reductions. Manholes, loading and unloading equipment, and excess doors for handling catalyst modules are required in larger, stationary SCR applications as well. Most custom-engineered SCR systems have space available for adding a row or layer of catalyst at a later date to reduce the emission even further or to meet certain catalyst replacement strategies. (See Fig. 11.)

The SCR Catalyst This is the ‘‘heart’’ of SCR systems and much research and development work went into catalyst material formulations, structures, and production processes over the last 20 years. Initially, catalyst ‘‘fouling’’ was a major problem, with the following concerns:

• • •

Masking, washable flue gas deposits reduce the reactive catalyst surface temporarily Poisoning, an irreversible degradation of the catalyst surface Plugging, dust clogging of the catalyst, causing an increase in back pressure and/or reducing catalyst reactivity/ performance.

The operating temperature window was limited to 570–850°F (300–450°C) to avoid ammonia sulfate formations at low reactivity and ammonia combustion problems at high temperature. These problems gave the SCR technology in the early developments prior to the 1990s a bad name. Since then major R&D and

106


FIG. 11 Reactor housing with SCR catalyst in shelf system. (Courtesy of EESI/Steuler.)

application engineering work was carried out in Europe. Also, a pollution control industry, trucking industry and air quality regulator partnership conducted lab/bench and field tests with several technical universities and certified test labs in Europe. Various catalysts from different European and U.S. manufacturers were evaluated in late 1980s to early 1990s Some test results from the technical university RWTH-Aachen, Germany of 1989–1991 are shown in Table 1 and Fig. 12. Upon the completion of these tests, the best performing ‘‘Catalyst A,’’ the diesel catalyst from Siemens was selected by the European trucking industry for the field tests in the early 1990s. Today, several different types of SCR catalyst are commercially available from several suppliers. The diesel SCR catalyst did overcome past SCR problems related to high-sulfur fuel, a phosphorus and zinc compound containing lubricating oils, arsenic resistance, and high-dust loads and allows operating temperatures as low as 300°F (150°C) and as high as 1020°F (550°C). Like most oxidation catalyst guaranties today, the SCR catalyst service life was guaranteed for 1 year only during early developments. Today, the standard process or performance guarantee for SCR systems worldwide is 20,000 operating hours or 3 years, whichever occurs first, in stationary and several hundred thousand kilometers in mobile, on-road applications, maintaining the emission reduction, ammonia slip, and so forth.

Tank and Piping System for the Reducing Agent Anhydrous Ammonia, Aqueous Ammonia, or Aqueous Urea Anhydrous ammonia, a hazardous, toxic, and flammable gas, requires professional handling. The operator of supply piping and anhydrous ammonia storage tanks is

Homogeneous, extruded Coated substrate (CS-based) CS based CS based Homogeneous extruded zeolite CS based

V2O5 /TiO2 /WO3 V2O5 /TiO2 V2O5 /TiO2 V2O5 /TiO2 /WO3 Fe/molecular sieve Cu/zeolite

European HDD Truck Tests, Catalyst Selection

45 100 100 100 21 100

Cpi 5100 8200 8200 8200 4000 8200

SV(m3 /hm3)

Note: Cpi ⫽ Channels per inch, SV ⫽ space velocity, AV ⫽ area velocity, A/V ⫽ area/volume.

A. B. C. D. E. F.

TABLE 1

6.0 5.8 6.8 6.8 7.0 6.8

AV(m3 /hm2)

850 1400 1250 1250 570 1250

A/V(m2 /m3)

Selective Catalytic Reduction 107

108


(a)

(b)

FIG. 12 NOx (a) and VOC/HC (b) reduction rates at different temperatures during 30–50 operating hours with NH3 slip of ⬍30 ppm.

required to strictly adhere to safety instruction of the manufacturers as well as to local, state, and federal regulations. Aqueous ammonia is safer but requires the same handling and storage precautions as for anhydrous ammonia. In Europe, 25% ammonia concentration in demineralized water is standard. In the United States, sometimes 27–29% ammonia concentrations are common, requiring pressurized tank certifications and routine maintenance in some states (i.e., New Jersey). Ammonia with a concentration of 20%, however, usually does not require such certified tanks. Engineers with little or no corrosion engineering experience sometimes specify carbon steel rather than SS316 for piping and lower-grade stainless steel for storage tanks in case aqueous ammonia is used, not realizing that aqueous ammonia is highly corrosive in the vapor phase. This tank corrosion has been causing extensive downtime due to rust particles, clogging valves, spray nozzles, and so forth. Again, this gave the SCR technology a bad name in the United States. Examples of SCR systems with anhydrous and aqueous ammonia are shown in Fig. 13. Aqueous urea, as salt dissolved in demineralized water, is nonhazardous, nontoxic, and nonflammable and is primarily used as fertilizer and animal feed. It required extensive R&D and application engineering development to make urea SCR systems work properly and reliably. The urea SCR system engineering development work in Europe was a prerequisite for applying the SCR technology to mobile, on-road, and off/nonroad applications, now benefiting the new and retrofit stationary combustion applications as well. Data on aqueous urea is presented in Fig. 14.

The Ammonia and Urea Supply, Metering, and Injection Systems These systems are due to the corrosive nature of the reducing agents made of SS316 steel. To achieve proper mixing of the reducing agent into the flue gas, the reducing agent is heated and diluted with hot air or recycled exhaust gas prior to


109

(a) Anhydrous Ammonia Tank

(b) Aqueous Ammonia Tank FIG. 13

(a) Gas turbine application with PEMS and CEMS controls. (b) Gas/diesel (dual fuel) 4.8MW engine cogeneration plant. (Courtesy of EESI/Steuler.)

being injected into the exhaust gas stream in front of the catalyst bed. Alternately, aqueous ammonia and aqueous urea may be co-injected and atomized with compressed air directly into the exhaust pipe. Because the dissolved urea salt may crystallize again and deposit around the spray nozzles, temperature control and other design considerations have to be met. Also, aqueous urea requires a certain reaction time to convert to ammonia in the exhaust duct, determining the injection nozzle location upstream of the static mixer.

110

Selective Catalytic Reduction Aqueous Urea: (NH2)2 CO ⫹ H2O, industrial grade Safety Aspects: Virtually nonodorous, nonhazardous, and nonflammable Consumption: Molecular weight ratio NO2: Urea ⫽ 1:0.87 NO: Urea ⫽ 1:1.00 Concentration: Urea in demineralized water (%) 30 – 40% Crystallization: ⫺11.4°C or ⫹11.5 F at 32.5% concentration Salting-out temperature: ⫹ 0.6°C or ⫹33.0 F at 40% concentration Specification: Typical industrial grade Value Dimension Example, Urea 40 ⫾ 0.5 % pH 8.0–9.0 Density at 15°C 1105–1125 g/cm3 Viscosity (at 25°C) 1.38 mPa s Specific heat (at 25°C) 3.26 kJ/kg K m S/cm Electrical conductivity (1.2–1.3) ⫻ 10⫺4 Temperature range without 0–35°C (32–95 F) temperature control Biuret ⬍0.5 % Fe ⬍0.1 mg/kg ⬍0.5 mg/kg PO4 Mg ⬍0.1 mg/kg Ca ⬍0.1 mg/kg FIG. 14 Aqueous urea data. (Courtesy of Siemens.)

There are basically two types of supply, metering, and injection system designs currently being used. The first is a constant high-pressure common rail piping system with needle valves or unit electronic hydraulic injectors. The alternate system incorporates corrosion-resistant, variable speed or variable stroke metering pumps of different designs. The reducing agent is metered and injected continuously in accordance with the signals received from the microprocessor or Process Logic Controller (PLC)based SCR electronic operating control system. Such system may feature either a feed-forward PEMS-based injection control, operating on historical computer memory data, or the feed-forward control plus a slower, but fine-tuning feedback control. The feedback system incorporates either a Continuous Emission Monitoring System (CEMS) or sensors. Such custom-designed SCR systems used to occupy an entire row of control cabinets. The mobile SCR system developments reduced such controls to less than laptop computer size. (See Fig. 15.)

The Integrated, On-Line SCR Operating Control and Emission Monitoring System This features a Predictive Emission Monitoring System (PEMS), which operates on historical emission data measured at different loads of the combustion equipment during (for example, engine bench tests or gas turbine, engine, or boiler SCR system) start-ups. The correlation function of engine-load values versus NOx emission values generated is combined with an algorithm in the software of the SCR control system, calculating the amount of reducing agent required to reduce the NOx to the specified


111

(a)

(b) FIG. 15 (a) SCR control system cabinets with PEMP and CEMS. (Courtesy of EESI/Steuler.) (b) Laptop-size SCR operating and injection control system. (Courtesy of Siemens.)

permitted emission rate. An example of this ‘‘mapping’’ process is shown in Fig. 16a–16c. Figure 16a shows the highest temperature of 900–950 F at 1000 rpm and medium torque, a typical truck operating condition. Figure 16b shows that the highest NOx emission is generated at medium speed and torque as well.

112


(a)

(b)

(c)

(d)

FIG. 16 (a,b) PEMS mapping: correlation functions of uncontrolled NOx and operating temperatures of a 12-L HD diesel truck engine rated 400 HP. (c, d) PEMS mapping: correlation functions of NOx reduction rates and break-specific fuel consumption. (Courtesy of Siemens.)

Figures 16c and 16d show that the highest NOx reduction takes place at medium speed and medium torque (Fig. 16c), where the lowest fuel consumption is achieved (Fig. 16d) as well. Engine test standards should therefore emphasize those operating conditions. The control system monitors all SCR system functions such as tank level control with high, reordering, and low limit, operating parameters such as exhaust gas temperature and pressure drop of the SCR catalyst as well as all required maintenance and trouble-shooting management functions. Thereby, the system communicates with, for example, the electronic engine management system and the remote central control panel or On-Board Diagnostic (OBD) system by CAN bus or via 4–20-mA or 0–5 V analog signals. In many stationary applications, the local air quality regulator requires the use of certified CEMS in accordance with U.S. EPA regulations such as 40 CFR Part 60, Appendix B for IC engines. Thereby, the accurate but expensive gas-analyzerbased CEMS are often more capital-intensive than the SCR emission reduction equipment, achieving NOx reductions of up to 90%. Again, this gave the SCR technology a bad name.


113

Catalyst Selection Process The type of SCR application determines the material and component selection. Not all SCR catalysts and system components perform well in all SCR applications. The SCR catalyst selection begins with a careful analysis of the performance and guarantee requirements. Then, the SCR catalyst meeting the performance characteristics best is selected. In some cases, this is not sufficient and the SCR catalyst is optimized through changes in catalyst material formulations; for example, reducing the reactivity to avoid exhaust duct corrosion through SO2 ⬎ SO3 conversion (SO3 ⫹ H2O ⬎ H2SO4), or changing the catalyst structure such as the cell density of honeycomb catalyst from i.e. 200 to 300 cells/in.2, to reduce the weight and the size of the SCR catalyst for a mobile on-road application where it matters most. In several cases, oxidation catalyst manufacturers in the United States also included oxidation catalysts upstream of the (SCR) catalyst, causing high corrosion through up to 80% SO2 ⬎ SO3 conversion and reducing the NOx emission reduction by converting more than 50% of NO ⬎ NO2, as shown in Fig. 17a–c.

SCR Catalyst and System Performance Parameters Depending on the application, different operating conditions and performance requirements have to be met. The following is a partial list of parameters that may have to be considered while designing a SCR system:

• • • • • • • • • • • • • • •

Fuels and fuel analysis: No. 1, 2, 3, and 6 fuel oil, natural gas, digester or landfill gas, wood chips, chemical waste, liquid or gas, as well as the chemical analysis thereof Fuel operation: different fuels as percentage of total operating time Flue gas analysis: NO/NO2 ratio, uncontrolled NOx and emission such as CO, VOC/HC, PM and NH3-slip as well as the emission reduction requirements NOx —in: raw/uncontrolled emission at different loads of the combustion equipment NOx —out: reduced, permitted NOx rate [kg or lb. per hour, ppm vd (dry by volume)] at stack or tail pipe Exhaust gas mass flow rate (kg or lb. per hour, N m3 per hour or scfm) and gas density Flow rate (m/s) and uneven gas flow and temperature deviations (%) at front face of catalyst bed, requiring computer and/or scale-flow-modeling and static mixer applications Maximum allowable pressure drop of the catalyst bed Pressure drop, total from combustion equipment to stack or tail pipe exit Sound attenuation requirements (dB A) Free-flow, area-, and space-velocity data Catalyst space availability: maximum cross section and length Temperature range, max. and min. at front face of catalyst at different loads H2O concentration in the flue gas O2 concentration in flue gas

114


(a)

(b)

(c) FIG. 17 (a) Metal-substrate based noble metal oxidation catalysts. (Courtesy of Miratech/Hug.) (b) Catalytic reactivity of noble metal catalyst for SO2 ⬎ SO3. (Courtesy of Dudoco 1995.) (c) NOx reduction rate at increased pre-oxidized NO ⬎ NO2 at ⬃500 F (250°C). (VDI-Report 1995.)


• • • • • •

115

O2 concentration of air quality regulator’s standard for emission rate calculations at 5%, 7%, or 15% O2 Acceptable process guaranties and equipment warranties Calculated catalyst bed dimensions and number of catalyst modules Oxidation catalyst requirements Particulate filter requirement Other

There are several SCR and Oxidation Catalyst and Application Engineering Calculations required to layout a SCR system. For some basic calculations equations are provided in Fig. 18.

SCR Catalyst Material There are basically three types of material used for SCR catalysts, noble, base metal, and zeolite. The material formulations and manufacturing processes are usually proprietary developments of the manufacturers. The reaction of NOx with ammonia takes place at the catalyst macropore surface, which may amount to 60 m2 / per gram of material of noble or base metal catalyst materials. In the case of the ceramic-zeolite-type catalysts, the exothermic reaction takes place inside the vast micropore structure of over 200 m2 per gram of zeolite. Noble catalyst metals are platinum, rhodium, and palladium. They can be used for both NOx reductions and for the oxidation of VOC/HC, CO, PM. Due to the high cost, primarily oxides of base metal are being used for SCR catalysts. Noble metal oxidation catalysts may be used upstream and/or downstream of the SCR catalyst: for upstream, to enhance the SCR NOx reduction by partially oxidizing NO to NO2, which in some cases, however, is counterproductive if SO2 is converted to SO3 as well; for downstream, to reduce possible ammonia slip spikes and CO/ HC not oxidized by the SCR catalyst. In HD diesel engine SCR applications using the Siemens diesel SCR catalyst, oxidation catalysts are generally not recommended because additional PM would be generated when burning sulfur fuel. Also, today’s efficient HD diesel engines emit only minimal amounts of CO and VOC/HC. Base-metal-based SCR catalysts contain oxides of base metals such as titanium (TiO2), vanadium (V2O5), tungsten (WO3), and additive and ceramic binders. V2O5 is highly reactive and used in small amounts of up to ⬃2% only. Catalysts with a high V2O5 content are used in the production of sulfuric acid (H2SO4) as well, which would also form in exhaust gas ducts if SO2 oxidizes to SO3 catalytically (SO3 ⫹ H2O ⬎ H2SO4). Base-metal-type SCR catalysts have nondiscrete macropores and channels, adsorbing ammonia, which is desorbed in a subsequent operation. This allows the adsorption of unreacted ammonia spikes (ammonia slip) rather than passing through the stack as secondary emission. Ammonia slip rates as low as 3–10 ppm have been achieved in continuous operations. The advanced diesel SCR catalyst development allows NOx emission reductions at temperatures as low as 300 F (150°C). Although originally developed for NOx reduction only, the advanced SCR catalysts is able to simultaneously reduce VOC/HC by up to 95%, PM by up to 50% and NOx by up to 95% at no extra cost. In Fig. 21, the plate-type and the extruded-type base metal catalyst packaging are pictured.

116

Selective Catalytic Reduction (a) The K values, the reactivity value of catalysts, may vary greatly, depending on type, material, and structure of catalyst Keff. ⫽

SV ⫻ ln(1 ⫺ n), AV

(1)

where SV is the space velocity, the volume of the exhaust gas flow (N m3) per hour at normal conditions divided by the volume of the catalyst (m3) resulting in (1/h), AV is the area volume, the catalyst surface area per catalyst volume (m2 /pm3), ln is the natural logarithm, and n is the emission reduction rate (i.e. 95% ⫽ 0.95). (b) Catalyst Volume (Vcat.) Vcat. ⫽

VEGF (N m3 /h) (m3) SV (1/h)

SV ⫽ Keff. ⫽

(2)

AV see Eq. (1) ln(1 ⫺ n)

where VEGF is the total exhaust gas flow at standard or normal condition (N m3 /h) and SV is the space velocity. [For each proprietary catalyst formulation and structure variation, the manufacturer has developed proprietary space velocity table values for reactivity and NOx reduction rate (1/h).] (c) Exhaust Gas Flow, (VEGF) VEGF ⫽ VEGF min. ⫹ NGcons. (N m3 /h)

(3)

VEGF min. ⫽ VAIR min. ⫻ NGcons. (N m3 /h)

(4)

VAIR min. ⫽ l⫽

2(Cx ⫻ Hx) (N m3 /h) 21

(5)

21 (%) (21 ⫺ O2act.)

(6)

where VEGF min. is the volume of air consumption times O2 content times NG consumption (N m3 /h), VAIR min. is the volume of air with 21% O2 required to oxidize total HC (N m3 /h), l is the percentage O2 in the air used during combustion, Cx Hy is the various hydrocarbons (HC) of the NG analysis making up the total HC, O2 act. is the actual oxygen (O2) concentration of the exhaust gas, and NGcons. is the natural gas consumption (N m3 /h). (d) Reducing Agent Consumption, Example Aqueous NH3 Consumption NH3 cons. ⫽ VEGF(N m3) ⫻ NOx reduction (ppm) ⫻ 3.3/1,000,000 (kg/h) using aqueous

(7)

ammonia with a 25% ammonia concentration, NOx —Reduction ⫽ (NOx in ⫺ NOx out)(ppm),

(8)

ppm NO ⫽ ppm NO ⫻ 30 (molecular weight)/22.4

(mg/N m )

ppm NO2 ⫽ ppm NO2 ⫻ 46 (molecular weight)/22.4

(mg/N m3),

3

ppm NH3 ⫽ ppm NH3 ⫻ 17 (molecular weight)/22.4

(mg/N m ),

ppm SO2 ⫽ ppm SO2 ⫻ 64 (molecular weight)/22.4

(mg/N m3),

O2act vs. stand ⫽ (ppmvd compound)

(9)

3

21 ⫺ O2 stand. (ppm vd) 21 ⫺ O2 actual.

(10)

FIG. 18 (a) Reactivity of SCR catalysts; (b) catalyst volume calculation; (c) Exhaust gas flow calculations, natural gas combustion; (d) reducing agent consumption.


117

FIG. 19 Pellet-type catalyst. (Courtesy of SCAQMD.)

Zeolite, also called molecular sieve, is a ceramic material. Some zeolite structures occur naturally; others like the ZSM5 of Mobile Oil is produce synthetically. Oil refineries heavily depend on them for their gasoline cracker, synthetic lubrication oil, and gasoline from natural gas processes. The extruded honeycomb-type zeolite-based SCR catalyst has a very large micropore structure of over 2000 ft2 or 200 m2 per gram of material. NOx and NH3 are attached to the micropore surface ˚ in size. This enorupon passing through the discrete pore openings of ⬃6–10 A mous sponge effect compensates for major spikes of NH3 and NOx during rapid load changes. The exothermic reaction of ammonia and NOx takes place inside the micropore structure through electrostatic forces. The reaction products, N2 and H2O vapor, are disposed of, back into the exhaust gas. This reaction is relatively slow, requiring a higher volume of catalyst than, for example, base metal catalysts. However, the zeolite catalyst has superior resistance to many compounds such as heavy metals, which are unable to enter the micropore structure through the discrete openings, and thereby extending the service life of the SCR catalyst considerably. One example of the over 1000 known different zeolite crystals is shown in Fig. 20. Some other more novel combinations of zeolite with noble or base metal materials are presently being researched for PM and other emission reductions.

SCR Catalyst Structure There are three types of SCR catalyst structure. The pellet-type catalyst, the extruded monolithic, honeycomb-type catalyst using either oxides of base metals or zeolite, and the coated-substrate-type catalyst, incorporating either corrugated foil

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FIG. 20 A zeolite crystal.

or plate-type stainless-steel sheet metal or extruded Corderite ceramic substrates. There is also a novel catalyst development incorporating fiber-based substrates. (See Figs. 20–22). Pellet-type catalysts are filled in containers through which exhaust gas is passed. The pulsing exhaust gas flow, however, cause the pellets to vibrate, abrade/ erode, and dust. The dust settles, clogging the catalyst bed and prevents an even gas flow. Due to the erosion of the pellets, the catalyst bed shrinks and unreacted exhaust gas will bypass together with the injected ammonia over the top of the catalyst bed into the atmosphere. Thus, pellet catalysts do not work most of the time and were replaced as soon as the honeycomb-type catalysts became available. However, there are still several such reactors operating in southern California today.

FIG. 21 Left: Coated stainless-steel mesh/expanded metal-substrate-based plate-type catalyst for high-dust applications; right: extruded, honeycomb, monolithic, base metal diesel catalyst. (Courtesy of Siemens.)


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FIG. 22 Macropore structure of extruded base metal catalyst.

The extruded, monolithic, honeycomb-type SCR catalyst has low back pressure and is widely used for gas turbines and boilers, engines, and other applications. The higher the number of extruded channels per square inch (cps), the higher the reactivity/active surface area and the smaller the catalyst for a specific application. Advanced base-metal-type SCR catalysts are available with 14–300 cps with channel wall thicknesses of 0.3–1.08 mm containing the macropore structure. This advanced catalyst development allows NOx emission reductions at temperatures as low as 300 F (150°C). The lack of a ‘‘sponge effect’’ may also be the reason for the lower emission reduction rates achieved by the coated-substrate-based catalyst at temperatures below 480 F (250°C); see Fig. 23. The coated-type SCR and oxidation catalyst has usually three layers: the corrosion-resistant substrate (such as the extruded Corderite monolith, corrugated stainless-steel foil or mesh plates, the aluminized washcoat to which the third layer, the catalytically active material, is bond. The corrugated foil substrate is primarily used for noble metal catalysts, whereas the Corderite monolith is used for noble and base metal. The plate-type catalyst has been developed for flue gases, containing high-dust loads, such as the hard coal utility boiler, industrial and municipal solid-waste incineration, and other industrial applications. Long-term operating experiences in Europe showed that the erosion of the reactive catalyst material at the face of the SCR catalyst bed will terminate upon the exposure of the stainlesssteel substrate, extending the service life. Due to the smaller macropore structure and surface area and thus absorption capability, the coated-type catalyst is less reactive.

Conclusion It would be beyond the scope of this introduction to the SCR technology to go into further details of the process and the application engineering (i.e., review basically 10–20-year-old designs for coal and gas utility boilers). The future of

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(a)

(b) FIG. 23 (a) Extruded monolithic honeycomb catalyst macropore structure versus the reduced coatedmetal substrate-based macropore structure. (b) Standardized test: relative reactivity of extruded (⫽1) versus coated-catalyst structures depending on thickness of coating (⫽0.6 and 0.2) at space velocity of 60,000/h. (Courtesy of Siemens.)

the SCR technology lies in distributed power generation applications such as generation sets, cogeneration sets, and mobile on-road and non-road applications. There are already close to 1000 IC engine and turbine applications in service worldwide today (Intermacom AG). This number could multiply when HDD trucks and other mobile SCR applications come to market in 2001 through 2010. In the following section, a few examples of SCR projects are summarized. However, because some past design, application engineering, and operation deficiencies gave the SCR technology a bad name in the United States, SCR systems engineers will have to pay more attention to design and application engineering details in future (Table 2).


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TABLE 2 Why Certain SCR Systems Have Not Performed in the United States Failure Catalyst 1. Clogging and bypass of NOx and NH3 slip 2. Reactivity loss

3. Emission spikes at rapid load changes

System Design 1. Clogging of injection system valves or nozzles 2. Clogging of catalyst

3. Low emission reduction

Controls and Other 1. High emission spikes at load changes

2. Not cost effective

3. Politics: Operator’s good references but bad performance of air pollution control equipment in actual operation

Cause Pellet type catalyst Masking, poisoning, or delamination of catalyst coating with metal or Corderite-based substrates Little to no adsorption/ desorption capability of catalyst (nonmonolithic catalysts) at ⬍500 F (250°C) Heavy corrosion/particle volume due to carbon steel aqueous ammonia tank and piping material Carbon steel reactor housing, scaling/ particles due to temperature cycling Uneven gas flow at front face of catalyst bed or insufficient mixing of exhaust gas/NH3

Relying only on a downstream CEMS with long feedback/response time Including a fully certified CEMS, which is often more costly than the SCR emission reduction system for NOx, VOC, and PM itself Operators avoiding fines, shutdowns, and lawsuits of poorly maintained system or new/unproven technology, ‘‘promoted’’ by the regulator

Advanced (SCR) Technology Solutions Honeycomb or plate-type catalyst Special lean burn/diesel catalyst, homogeneous material, allowing up to 3% sulfur fuel and an operating temperature window of 300–1020 F Diesel catalyst with micropore structure/‘‘sponge effect’’ with adsorption/ desorption features

All stainless-steel storage, delivery and injection system Heat-resistant steel such as low Molly steel

Gas flow modeling, scale model tests, and low back-pressure static mixers such as Parmix TM/ TM Siemens Feed-forward PEMS-based control with optional feed back CEMS or sensor-based control Electric–chemical sensorbased accurate spot check analyzer with periodic emission testing by third party Independent test lab certification, confirming equipment manufacturer’s long-term performance claims during a 3-year performance guarantee

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Summary of SCR Application Case Studies Diesel/Coal-Slurry Fueled Diesel Engine Cogeneration Plant, University of Alaska, 1999

System Overview The University of Alaska, Fairbanks (UAF) has a 9000-kW electric Fairbanks– Morse reciprocating engine, which currently is fueled on No. 2 diesel fuel upon start-up in Summer 2000. UAF, in partnership with Fairbanks–Morse and the Department of Energy (DOE), has been constructing a demonstration project to test the feasibility of pulverized coal mixed in water, a coal slurry as an alternative fuel. Fairbanks–Morse had already developed and tested the engine modifications in their research facility in Wisconsin so that the long-term field tests could commence in 2000. The engine will provide electrical power for the campus. Power not used by the campus is sold to the local electric utility. An exhaust waste heat recovery system generates steam, which is used for heating and cooling campuswide. An SCR system was proposed to reduce NOx emission. Because of the relatively high sulfur content of the fuel and the required low SO2 to SO3 conversion, a zeolite catalyst was selected. When operating on coal slurry, the engine produces large amounts of particulate matter (PM). In order to deal with the high-PM load, the extruded monolithic honeycomb SCR catalyst was designed with a larger than normal pitch or channels per square inch (CPSI). This allows the PM to pass through more easily. The system was designed with a vertical exhaust flow, from top to bottom, allowing the PM to pass through and being collected in an ash hopper below the catalyst housing. A soot-blowing system has been installed above the catalyst. The soot blower automatically blows down the PM from the catalyst at regular intervals when the engine is running on coal slurry. When running on diesel, the exhaust gas PM or soot concentration is low enough to operate without the soot blower. The SCR system is designed to reduce NOx emissions independently of the fuel used. Because there is a large difference in NOx emission, exhaust temperature, and exhaust flow rate between the two fuels, the system had to be designed with a wide operating range. When running on diesel fuel, exhaust temperature and flow, and NOx emissions are higher than when running on coal slurry. A system of duplex metering pumps was furnished. The reducing agent flow is adjusting by varying the speed of the metering pumps. When operating on coal slurry, only one pump is required. When operating on diesel, both pumps will run. The system is designed to use aqueous ammonia (20% NH3 in demineralized water by weight) as the reducing agent. Ammonia was selected over urea because it is easier to source in the region, and because urea would require a great deal of temperature control during winter operation. Injection rate control utilizes a feedforward system, which sets pump speed based on engine load. A control feedback is used which measures NOx emissions upstream and downstream of the catalyst; this data are then used to ‘‘trim’’ or fine-tune the reducing agent injection. The continuous emission monitoring system (CEMS) analyzer (furnished by UAF) alternately reads emissions upstream and downstream of the catalyst and thereby


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provides the feedback necessary to ensure that the SCR system is operating at peak emission reduction with minimal ammonia slip. The system is designed to reduce NOx emissions by 90%, with ammonia slip of 10 ppm or less. Currently, since the system start-up in Summer 2000, the system is operating within these parameters, on diesel fuel. During commissioning, a 90% NOx reduction and almost zero ammonia slip was measured. The construction on the coal–water slurry production facilities will start in 2001. The engine is scheduled to run primarily on coal slurry by late 2001 or early 2002. In Fig. 24, the process and instrumentation diagram (P&ID) and modified summary excerpts of the operating and maintenance manual are presented. Figure 25 shows the Table of Contents of the summary excerpt of the Operating and Maintenance Manual.

Design Parameters The SCR system is designed based on the parameters shown in Fig. 26.

SCR Catalyst and Housing CER-NOx (SCR) Abatement Catalyst. The CER–NOx* (SCR) NOx abatement catalyst is a honeycomb-type, molecular sieve, all-zeolite catalyst. Zeolites ˚ . The are crystalline microporous ceramic solids with pore openings of 3–10 A 2 micropores provide over 1500 ft of surface area per gram of zeolite material. The SCR catalyst is located in the exhaust gas stream downstream of the Fairbanks–Morse engine, before the heat recovery boiler. Each catalyst module has dimensions of ⬃152 mm ⫻ 152 mm ⫻ 1000 mm long, without wrapping. Each module has openings with a 6-mm pitch/cps (channels per square inch) to allow particulate matters (PM) to pass through without plugging. Seventy-two modules are wrapped in a stainless-steel cartridge, which protects them from mechanical damage. The catalyst cartridge weight is 1510 kg (3329 lbs.) and its dimensions are 1920 mm ⫻ 960 mm ⫻ 1365 mm height. Chemical Process. Ammonia reacts with NO and NO2 within the zeolite catalyst micropore structure to form nitrogen gas (N2) and water vapor (H2O). Nitrogen oxides and the injected aqueous ammonia are removed from the exhaust gas through adsorption into the catalyst micro pores of the zeolite, based on the concentration gradient. Electrostatic forces generated inside the micropores decrease the activation energy for the reduction process, thus allowing reactions to occur in a temperature range of 300° C/570° F to 480° C/900° F. The reaction releases energy, which forcibly expels the reaction products N2 and H2O from the micropores.

* CER-NOx is a trademark of EEST/Steuler.

FIG. 24

Process and instrumentation diagram (P&ID) of the Alaska cogeneration plant. (Courtesy of EESI/Steuler.)

124 Selective Catalytic Reduction


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CER-NOx (SCR) NOx ABATEMENT SYSTEM Overview and Component Description Table of Contents 1. Design Parameters 2. (SCR) Catalyst & Housing a) CER-NOx (SCR) Abatement Catalyst b) Chemical Process c) (SCR) Reactor Housing 3. Aqueous Ammonia Supply & Injection System a) General Description b) Ammonia Metering Panel c) Compressed Air Subsystem d) Electric Sub-panel e) Ammonia Injection Lance Assembly f) Static Mixer g) Ammonia Pump Station h) Ammonia Storage Tank i) Ammonia Tank-Truck Unloading Station 4. Operating System Control a) Condition b) Analysis c) Action 5. Operating Routines a) Start Up b) Normal Operation c) Shut Down 6. Maintenance Routines 7. Daily Log, Visual Inspection FIG. 25 Table of Contents of the CER–NOx SCR system.

SCR Reactor Housing. The housing is fabricated of A36 carbon steel. It includes the following:

• • • •

A 90° inlet transition from the exhaust duct to a turning vane assembly to distribute the exhaust gas evenly across the catalyst bed A shelf system with one row for catalyst One soot-blowing system to blow ash off the catalyst bed during coal–water fuel operation. The system includes control valves, isolation valves, and drive motors, controlled from the Co-gen plant control system Bolt-on access doors provided in order to load and unload catalyst; catalyst is loaded into the housing after the housing has been installed

Aqueous Ammonia Supply and Injection System General Description. The reducing agent supply and injection system consists of a storage tank, an ammonia pump station, a metering panel, an injection spray nozzle, and interconnecting piping. A static mixer is welded into the inlet of the reactor housing to properly mix the reducing agent and to evenly distribute the

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• •

Emission source: Fairbanks–Morse dual-fuel engine, 9000 kW(e)/ ⬃12,500 BHP Fuel type: either diesel fuel or coal–water slurry fuel

Combustion Process Information: Diesel Fuel

Coal Slurry Fuel

Exhaust flow rate, lb/hr Exhaust temp., max, °F Exhaust temp., design, °F

155,800 837 665

105,400 800 765

CO2, weight % O2, weight % H2O, weight % N2, weight % NOx, volume SO2

8.6 13.3 3.5 74.4 0.18 vol.%, wet 0.016 vol. %, wet

11.3 11.9 6.6 70.2 600 ppmv 79 ppmv

Ash/particulate, lb/hr

2.2

350

—SCR System Performance NOx Reduction, % Ammonia slip SO2 to SO3 conversion Pressure drop, max., for clean system —Reducing Agent —Reducing Agent consumption —Atomizing Air Requirements Pressure Consumption

90% — 10 ppmvd (15% O2) 0.1% max 1.75″ water column, measured from inlet hood to end of catalyst bed Aqueous ammonia, nominal 25% in water Technical grade in demineralized water only! 53.1 gal/h 21 gal/h Minimum 70 psi, maximum 120 psi 32 SCFM

FIG. 26 Design parameters for the SCR system. (Courtesy of EESI/Steuler.)

exhaust gas across the front face of the catalyst. The metering panel is controlled by the Co-gen plant’s Distributed Control System (DCS). The reducing agent is supplied via stainless-steel pipe to the metering panel. An ammonia pump station provides pressurized (⬃40 psi) ammonia to the ring line. The metering panel contains metering pumps, which controls how much reducing agent is injected into the gas stream. Refer also to Fig. 24. Ammonia Metering Panel. The metering panel is divided into two subsystems: one for reducing agent control, the other for control of atomizing compressed air. The reducing agent subsystem consists of the following components:

• • •

One fine filter (7 µm) on the pump outlet. Two metering pumps (P-103, P-104). During coal–water fuel operation, only P-103 is operated. When the engine runs on distillate fuel, P-103 is run at full speed and P-104 is operated at varying speed to provide the balance of the reducing agent volume required. Two variable-speed drives (SIC-103, SIC-104), one for each metering pump motor. Each drive is controlled by the DCS via a 4–20-mA signal (for speed setting) and dry contacts (for starting/stopping the drive). A 240-VAC single-


• • • • •

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phase input to each drive is converted to a 240-VAC three-phase, variablefrequency output to the metering pump motor. The drives are located in an electrical subpanel mounted on the side of the metering panel. One normally closed (energized to open) solenoid shutoff valve (FV-101). This valve gives a positive shutoff of the reducing agent feed when the system is turned off. The valve is controlled by the DCS. The valve is opened when a metering pump is started and closed when the metering pump is stopped. One pressure gauge (PI-105) on the pump discharge to monitor injection pressure. Two relief valves (PSV-103, 104) to protect the metering pumps. Two pulsation dampeners to absorb pulsations from the diaphragm pumps. One magnetic-inductive flow meter (FE/FIT-101) which provides feedback to the DCS to close the feed rate control loop.

Compressed Air Subsystem. This subsystem consists of the following components:

• • • •

One pressure regulator (PCV-201), to maintain air pressure at 43 psig. One filter/dryer, to remove particulate matter and moisture from the compressed air. One normally closed (energized to open) solenoid valve (FV-201). This valve closes off the airflow when the system is not running. It is controlled automatically by the DCS. One pressure switch (PSL-201). If the air supply pressure drops below 50 psig, the pressure switch contacts open, signaling the DCS.

Electrical Subpanel, Mounted on the Metering Panel. This subpanel consists of the following components:

• • • • • •

One 20-A disconnect switch, to provide a local disconnect of the 240-V power supply to the metering panel One 24-VDC power supply to provide power for the solenoid valves and flow meter Three circuit breakers to distribute AC power to the two variable speed drives (SIC-103 and SIC-104) and 24-VDC power supply One circuit breaker to switch on the outlet of the 24-VDC power supply Two loop isolators for the variable-speed drives (SIC-103 and SIC-104) speed setting input Four 2PDT pilot relays with 24-VDC coils, to take control inputs from the DCS for SIC start/stop control.

Ammonia Injection Lance Assembly. The injection lance is inserted into the exhaust duct upstream of the first static mixer. The lance’s spray head shall point in direction of flow. The lance consists of an air/liquid atomizing nozzle and a carrier pipe, which transfers the atomized reducing agent into the exhaust duct.

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Selective Catalytic Reduction Static Mixer. One static mixer is provided, which is welded into the duct upstream of the reactor-housing inlet. The mixer ensures proper ammonia/exhaust gas mixing. Ammonia Pump Station. The ammonia pump station supplies aqueous ammonia to the ring line feeding the metering pumps. It consists of the following:

• • • • • • • •

One supply pump, P-102, diaphragm type with fixed speed and capacity of 90 gal/h. One motor, M-102, 1/3 hp, 115-VAC single phase, TEFC to drive P-102. One pulsation dampener, to remove pulsations from the pump outlet. One inlet strainer, 190 µm, to remove larger particle from the aqueous ammonia. One pressure relief valve, PSV-101, to protect the pump. One pressure/vacuum gauge (PI-103) on the inlet and one pressure gauge (PI104) on the pump outlet. One back-pressure control valve (PCV-102) to set ring line pressure; the return side of the ring line is connected to the PCV. One control panel, with HOA switch, main disconnect, motor contactor, in NEMA 4 enclosure. This panel also houses the ammonia tank level indicator/ switch (LIS-101) which indicates the tank level in gallons.

Ammonia Storage Tank. The aqueous ammonia storage tank holds 8000 US gals, is made of 304 stainless steel, and is resistant to the highly corrosive ammonia vapor above the liquid. The tank is designed for atmospheric pressure. It is supplied with ball valves, a safety valve, and a level transmitter, shown on Fig. 24. Ammonia Tank-Truck Unloading Station. The unloading station is used to transfer aqueous ammonia from the delivery tank truck to the bulk storage tank. It consists of the following:

• • • • • •

Centrifugal transfer pump, P-101, with capacity of 100 gal/min. A 3 HP, 3450 rpm, 460VAC motor, M-101. Flow switch, FE/FS-101, interlocked to the motor, to shut down the pump if flow is lost. One pressure/vacuum gauge (PI-101) on the inlet and a pressure gauge (PI102) on the pump outlet. Cam-and-groove fittings for the tanker to connect. Control panel, NEMA 4 enclosure, with manual start/stop control switch and interlock with the flow switch above and the ammonia tank level switch (LIS101). Upon a ‘‘high tank level’’ alarm, the unloading pump will be shut down.

Operating System Control The SCR system is controlled by the central DCS of the Co-gen plant. The general operating sequence should follow the program in Table 3. See the detailed program description for pin assignments and input/output requirements.


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TABLE 3 Operating System Control Condition On start of engine

Advice from Engine control if running on distillate or coal–water fuel Exhaust gas temperature (570 F) (from TIT-301)

Exhaust gas temperature (896 F) (from TIT-301) Ammonia storage tank low level

Analysis

Actions

Begin purge of injection lance Start pump P-102 DCS decision on which metering pump to run

Send ‘‘start’’ signal to P102

Exhaust temperature for SCR Process OK, ammonia injection permitted

Open ammonia valve FV101 Start metering pump P-103 (if CWF) and P-104 (if distillate fuel) Stop metering pumps Close ammonia valve FV101 Stop metering pumps; close ammonia valve FV-101 Stop metering pumps; close ammonia valve FV-101 Stop metering pumps

Exhaust temperature for SCR Catalyst too high The ammonia tank empty

Atomizing air pressure low

Failure of compressed air system

Ammonia flow below min value (from FIT-101) and metering pump start signal is given, and metering pump speed signal above min value On shutdown of engine

Failure of metering pump; allow 1 min time delay from start of pump before writing alarm

Stop metering pumps; close ammonia valve FV-101; close air valve FV-201; stop ammonia pump P-102

Source: EESI/Steuler.

Operating Routines Start-up. Upon a complete equipment shutdown, the following restart checks have to be performed: No maintenance is being performed at the NOx abatement system, a visual inspection shows no disconnected or broken pipes, wires, and so forth, and the reducing agent supply is available.

• • • • • •

Open all hand valves in the reducing agent supply and injection lines Atomizing compressed air supply (compressor) available Open all hand valves of the atomizing compressed air supply and aqueous ammonia injection lines Verify that the DCS available Verify power is available to metering panel Start the system via the DCS upon determine permissive conditions to start

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Selective Catalytic Reduction injection based on temperature, lack of alarm signals, and other operating parameters Normal Operation. Under normal operating conditions, the DCS controls the speed of the reducing agent pumps in accordance with the NOx signal received from the continuous emissions monitor-based feedback. It is generally not required to make adjustments or continuously monitor the system. Shutdown. If the system will be shut down for maintenance, and so forth, the following steps have to be taken:

• • •

Shut down aqueous ammonia injection and allow system to purge lines and lances with compressed air Shut off power supply to the metering panel Close air and ammonia hand valves at inlet and outlet of metering panel

Maintenance Routines In general, routine maintenance on the CER-NOx SCR system is limited to occasional visual inspections, pump oil changes, and periodic replacement or cleaning of filters. Usually, the system can be quickly checked out in conjunction with other maintenance activities. Table 4 outlines how often certain maintenance procedures should be performed. Note that these intervals assume continuous operation. Under intermittent operation, these intervals could be extended. The operator shall determine if the intervals can be lengthened based on operating experience.

Daily Log Table 5 is the inspection checklist.

Gas Turbine Combined Cycle Power Plant Rated 2 ⫻ 170 MW, Natural Gas/Liquid Fuel Fired at Bridgeport Harbor, Connecticut

Overview Cogeneration and distributed power generation will be the preferred and most fuelefficient way of the future to generate heat and electricity as fuel cost increases and the utility industry becomes fully deregulated. Rather than ‘‘destroying’’ the so-called ‘‘waste heat’’ in cooling systems, the local heat requirement will be the driving output/product and the by-product, the excess electricity can easily be wired away. Fuel efficiency has been a driving force of air-pollution control regulator outside the United States in the power generation community, especially in Japan and Europe. Fuel efficiencies of over 85% versus 50–60% max. in


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TABLE 4 Maintenance Routine Operation Visual inspection (see: Visual Inspection Daily Log) Check reducing agent supply tank level Check/replace metering pump (P-103/ 104) and ammonia transfer pump (P-102) oil. Replace every 1000 h of operation Drain differential pressure gauge (PDIT-301) condensate lines for SCR reactor Record differential pressure indication for SCR reactor Clean/replace reducing agent in metering panel Clean atomizing air filter in metering panel Clean reducing agent strainer at ammonia pump station Check/replace metering pump and transfer pump diaphragms Check metering panel and pump station pulsation dampener charge Check/rebuild metering panel solenoid valves (FV-101/201) Check metering panel pressure relief valve setting (PSV-103/104) Check air pressure switch (PSL-201) setting Check injection lance nozzle for cleanliness Calibrate all 4–20-mA loops to DCS

Daily Weekly Monthly

Biannual

Annual

X X X

X

X X X X X X X X X X X

Source: EESI/Steuler.

combined-cycle, gas/steam turbine projects have been achieved. A Swedish utility company operates a 1200-MW cogeneration power plant, providing heating to the town nearby. The Saarbruecken utility company in Germany, reducing CO2 emission by over 15% in the prior 10 years, primarily through cogeneration, received the U.N. Environmental Award at the summit of world leaders in Rio, Brazil in 1990. In the United States, utility companies have rarely opted for the highly fuelefficient cogeneration plant alternative yet. The new or upgraded power plants, however, usually incorporate gas turbines and downstream steam turbines, utilizing the steam generated from the ‘‘waste heat’’ of the gas turbine for electricity generation.

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Selective Catalytic Reduction TABLE 5 Inspection Checklist Inspection check list

Initial when done

1 Inspect reducing agent storage tank, and stainer for good physical condition. Verify no leaks exist. 2 Inspect supply line between reducing agent storage tank and metering panel. Verify no leaks exist and that isolation valves are fully open. 3 Inspect atomizing air source. Verify that no leaks exist in air supply lines between atomizing air source and metering panel, and that isolation valves are open, and correct pressure is available (75–120 psi 4 Check metering panel for proper position of its isolation valves (fully open), proper draining of condensate drains, that no physical damage exists (internal or external of enclosure), and that no air or reducing agent is leaking inside or outside of enclosure. If system is running, that it is making normal operation noises, and when finished; that the door is closed and latched. 5 Inspect air and reducing agent lines between metering panel and injection lance for good physical condition. Verify no leaks exist and that isolation valves are fully open. 6 Check that injection lance is in good physical condition (no bent or cracked hook-up ports). 7 Inspect insulation on exhaust duct and SCR housing for condition and proper installation. Verify that no obvious exhaust leaks exist (indicated by soot tracks at insulation joints). 8 Check thermocouple probe at inlet of SCR housing that no exhaust leaks exist and that termination cover is in place. 9 Inspect the SCR differential pressure gauge and its associated block and bleed valves. Verify that they are not damaged and are properly positioned (high- and low-pressure input valves open, bypass valve closed, condensate lines closed). 10 Verify that the DCS is operating normally and that no alarms are displayed.

Project Summary The (2) Siemens V 84.3A gas turbine project is rated 170 MW electric per gas turbine. The process and instrumentation diagram (P&ID) of the SINOx exhaust gas cleaning system is shown in Fig. 27. The SCR system design is based on the exhaust gas data listed in Figs. 28a and Fig. 28b. The system ensures State of Connecticut EPA compliance with the emission reduction of 91% to 4.5 ppmvd at 15% O2, dry basis for nitrogen oxides (NOx, as NO2). The SINOx (SCR) system features the use of either aqueous urea or aqueous ammonia as the reducing agent to meet future, more restricted safety regulations. The SINOx system was delivered as two preassembled units.


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FIG. 27 P&ID of a (2) gas turbine power generation plant.

A brief technical description of the systems function is provided in Fig. 29 and a system maintenance schedule is given in Fig. 30.

Mobile, Portable and Other Applications

Background Summary California ARB (Air Resources Board) Strategy for Additional Emission Reductions by 2007/2010. During the University of California–Irvine Technology Meeting on October 6, 1999, the Cal.-ARB presented its objectives to substantially reduce NOx, HC, and PM emissions. To achieve California SIP (U.S. EPA’s State Implementation Plan) goals by the year 2007/2010 emission of onroad and off/nonroad vehicles and equipment could be reduced by market incentive and monetary incentive programs. Retrofit emission reduction applications for diesel locomotives, diesel-powered coastal vessels and construction/mining equipment, portable generation sets, and various agriculture (i.e., irrigation/pump drives) and garden equipment should be prime targets. Summary of Recent Market Data on Retrofit Emission Reduction Target Markets. Over 50% of the total U.S. NOx emission inventory of ⬃25 million tons per year and close to 60% of PM emissions is generated by portable, on-road, and off-road vehicles and equipment.

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(a)

(b) FIG. 28

(a) Project data, operating conditions of the (2) gas turbines; (b) project data, system utilities.


135

FIG. 29 Brief description of the SINOx system process.

•

A total of 2.2 million tons/year of NOx from the United States representing 8.6% were assumed from Class 8 HD diesel trucks. However, due to excess ‘‘off-test-cycle’’ emissions of 15.758 million tons during 1988 through 1998, caused by 1.328 million Class 8 trucks, an additional 1.3 million tons of NOx per year had to be added, for which engine manufacturers were fined in a con-

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FIG. 30 System maintenance schedule.

•

sent decree. This increased HDD truck NOx emission resulted in over 3.5 million tons of NOx emission per year, equal to 12% of the total U.S. NOx inventory. The Consent Decree of the U.S. Justice Department and the U.S. EPA, the State of New York, and Cal.-ARB with the engine manufacturers did not incorporate any short-term remedies but only engine-rebuilt solutions for this large fleet of highly fuel-efficient HDD trucks. One reason, according to industry, is the unrealistic U.S. FTP (Federal Transient Protocol) test cycle with heavy emphasis on low torque/low rpm and high rpm, whereas most truck operations take


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FIG. 31 Product summary, SINOx SCR System. (From Intermacom AG, Feb. 20, 2000, Draft for the WebPages of Cal.ARB.)

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•

place in medium torque/rpm. In comparison, Japanese and European tests cycles emphasize medium torque/rpm, almost neglecting the low torque/low rpm operation. One of these European tests shall now complement the U.S. EPA FTP tests. New trucks however, are required to meet NOx ⫹ HC emission levels of 2.4 gr./BHP h by 2002/2004. Off-road NOx and PM emission from diesel-powered locomotives, ships, construction, mining, and agriculture equipment as well as portable equipment such as generation sets have not been a U.S. EPA priority even though they account for approximately one-quarter of the total U.S. NOx inventory. By 2010, U.S. diesel locomotives are required to meet 5.6 gr./BHP h. California has an earlier target date and fleet averaging provisions. Also, marine applications lag behind, even though less than 2 gr./BHP h NOx emission rates have been demonstrated in close to 100 marine diesel engine applications (with engines rated 300 to over 10,000 BHP) in Europe. Portable generation sets are not required to achieve better than 5.9 gr./BHP h NOx emission rates in California, whereas some of the same engines in trucks have to meet 2 gr./BHP hr NOx in 2002/ 2004. According to Cal.-ARB there are 72,064 portable and stationary diesel engines without emission controls in California, rated 110–600 BHP.

Reports on Technology Evaluations. In recent publications of UC-Davis’ ITS (University of California, Institute for Transportation Studies) and Diesel Fuel News, various emission control strategies for HD diesel engines were discussed. Two most promising retrofit technologies were identified achieving over 70–80% NOx and VOC/HC and substantial PM emission reduction: The UREA–SCR and the NOx Absorber Technology. All other retrofit/post treatment technologies are either years away from any commercialization or achieve only 20–40% NOx reduction rates.

•

•

The NOx Absorber Technology has been tested by Cummins, using the Euro-3 (13 Mode) Steady State Diesel Engine Tests and 5 ppm sulfur fuel, achieving 80% NOx reduction at a 8.5% fuel penalty. In a Marathon–Ashland Petroleum commentary to U.S. EPA, it is claimed that only a 20,000-mile service life with 5–15 ppm sulfur fuel could be expected. Currently, no 15 ppm sulfur fuel is commercially available on a large-scale basis in the United States. The U.S. EPA does expect that such 15 ppm sulfur fuel will be readily available in the United States prior to 2007. In addition, lube oils with phosphor and sulfur compounds have to be reformulated and tested to avoid loss in engine service life. The UREA SCR Technology for the simultaneous reduction of NOx (70–85%), VOC/HC/AirToxics (80–95%), and PM (up to 50%) has already been used in HDD truck field tests in Europe and the United States in the 1990s. The technology will be commercially available by 2001, has been field tested for 4 years, and demonstrate no fuel penalty. The service life is expected to be over 300,000 miles and the initial target price is estimated to be US$ 2000–3000 per HDD truck. The European truck manufacturers and Siemens pioneered this technology. By the end of 2001, Siemens will go into the SCR system production for Daimler–Chrysler and MAN’s new ultralow-emission HD diesel trucks. The volume-produced SINOx Systems can then be used for mobile, transportable,


139

(a)

(b) FIG. 32

(a) Stationary and mobile applications for the diesel SCR catalyst. (b) Simultaneous reduction of NOx, VOC/HC, and PM at stationary and mobile applications. (Courtesy of Siemens.)

and stationary engines rated 100–600 BHP. However, to use the prefabricated, off-the-shelf SINOx products as retrofits kits, engine model/application adaptations through local factory trained and licensed dealers are required. The SINOx SCR systems are used in coal and gas-fired utility boilers and gas turbines, stationary and portable generation sets, cogeneration, and various mobile, on-road and off/nonroad applications. Figure 32 shows the range of SINOx Applications.

Jet Fuel Pipeline Pump Station for the New York Area Airports, NJ This SCR application was discussed in two U.S. engineering journals in 1998. Three 3000-BHP gas engines with low NOx emission exceeded the plants VOC/ HC/AirToxics, CO, and NOx emission limits, set by the State of New Jersey. Figure 33a shows one of the three engine enclosures with the SCR system and exhaust stack located in front. In Fig. 33b, the engine data are listed. The system design aspects are listed in Fig. 33c. The process guaranteed emission limits are shown by Fig. 33d.

140


(a)

(b)

(c)

(d) FIG. 33 (a) Engine enclosure with SCR system and stack in front; (b) Data on the three gas engines of the jet fuel pump station; (c) SCR system design data; (d) Pump station permitting and process guarantee data. (Courtesy of Siemens.)


141

Coastal and Ocean-Going Vessels and Other Marine Applications Two of the vessels equipped with SINOx diesel SCR Systems achieving very low emission rates of equal or less than 1 gr/pBHP h are pictured in Fig. 34.

Heavy-Duty Diesel Truck Field Tests in Germany and the United States, 1995–1999 SINOx (SCR) System Overview. The SINOx after-treatment system is a fully developed and tested diesel exhaust emission reduction system. Figure 35 pictures the first HD diesel truck on the left with a round muffler, replaced by a round SINOx diesel SCR System in the early 1990s. The right photo shows the Daimler– Chrysler HD diesel truck barrel-type SCR reactor/muffler, going into production at the end of 2001. The UREA–SCR System developed by Siemens (Siemens Westinghouse) and the European trucking industry uses aqueous urea as a reducing agent for the SCR process to reduce NOx. The process also reduces VOC/HC, and PM at the same time. The manufacturer has performed laboratory and fleet tests of the SINOx system. Approximately 20 European Mercedes Benz (Daimler–Chrysler), IVECO, and MAN heavy-duty Class 8 trucks were tested in the field. In late 1998, some of these trucks had accumulated over 300,000 miles in common carrier operations. The urea infrastructure, the risks of tampering with the UREA (SCR) system, and other initial operating reliability and other concerns were resolved to the European trucking industry’s and air quality regulators’ satisfaction. The diesel fuel used during the tests had a sulfur content of max. 500 ppm (0.05%). Figures 36a and 36b show program objectives and the aqueous urea pump station, respectively.

FIG. 34 SCR marine applications with extremely low NOx emission. (Courtesy of Siemens.)

142


FIG. 35 Pilot-type and preproduction model SCR system for Class 8 HD diesel trucks. (Courtesy of Siemens.)

(a)

(b) FIG. 36 (a) SCR truck program objectives; (b) SCR aqueous urea pump station. (Courtesy of Siemens.)


143

The SCR system feasibility tests included bench tests using European Stationary Cycle (ESC) (⫽OICA) and European Transient Cycle (ETC), and the Federal Transient Protocol (U.S. FTP) Test Cycle in the United States. NOx, VOC/HC, and CO of the exhaust gas were measured upstream and downstream of the catalyst, using sample gas conditioning systems, heated sample gas lines, and various gas analyzers. The following analyzers were used: chemiluminescent detector (CLD) for NO and NO2, flame ionization detector (FID) for VOC/HC, a nondisperse infrared analyzer (NDIR) for CO and CO2, and a magnetopneumatic analyzer for O2. Fourier transformation infrared spectroscopy (FTIR) allowed simultaneous, realtime monitoring for multiple gas components. The FTIR was used downstream of

(a)

(b) FIG. 37 (a) European steady-state cycle (ESC) test results; (b) European transient cycle (ETC) test results. (Courtesy of Siemens.)

144

Selective Catalytic Reduction the catalyst for the exhaust gas and for detecting potential traces of secondary emissions, such as ammonia slip (NH3), laughing gas (N2O), cyanic acid (HCN), formaldehyde (CH2O), and so forth. Also, a microwave process analyzer was used to measure ammonia slip during steady-state tests and for periodic verification/ calibration tests of the FTIR analyzer. For analytical hydrocarbon tests, bag samples were taken and analyzed with gas chromatographs. Test results of an European HD diesel engine with a SINOx System are summarized and compared with emission standards, based on the European steady-state cycle ESC (OICA) and the European transient test cycle (ETC) as shown in Figs. 37a and 37b, respectively. The NOx emission reduction was 80% in steady-state

(a)

(b) FIG. 38 (a) Urea SCR steady-state/13-mode test results, DDC-S 60, 12 L/400 HP HDD truck; (b) Urea SCR U.S. EPA’s FTP test results, DDC-S 60 engine. (Courtesy of Siemens.)


145

and 80% in transient tests. Hydrocarbons are simultaneously reduced by over 80%. Ammonia slip is limited to single-digit ppm, and no other undesirable secondary emissions were detected with the sophisticated analyzer setup described above. This was confirmed in U.S. bench tests in Detroit in 1998, using the most popular Class 8 HD diesel truck engine in the United States, the Series 60 Detroit diesel engine, as shown in Figs. 38a and 38b. The emissions measured downstream of the SINOx diesel catalyst was approximately 50% lower than the next lower tier of the European emission limits for NOx, VOC/HC, and PM [18]; that is, the emission limits achieved in Europe and later in the United States already met the year 2005 European and the 2002/2004 U.S. EPA emission limits in 1999.

SINOx System Description The exhaust gas generated by the diesel engine is fed into the SINOx catalyst, which is integrated into the exhaust gas pipe system (Fig. 39). A certain amount of the reducing agent, the aqueous urea (30–40%) is injected upstream of the SCR catalyst. The reducing agent storage tank has a volume of ⬃5% of the fuel tank volume. The control system data bus provides the data link between the HDD engine’s control unit and the SINOx control unit. A flow control unit in the metering panel meters the amount of reducing agent. The SINOx system is activated by the signals ‘‘engine on’’ (provided by the Engine Management System) and exhaust gas ‘‘temperature above set point’’ (temperature sensor in exhaust pipe). The amount of reducing agent is metered and in accordance with the computer-memory-based predicted NOx emission and injected into the gas pipe upstream of the SINOx

FIG. 39 SCR system layout. (Courtesy of Siemens.)

146


(a)

(b) FIG. 40 Utility boiler application.


147

catalyst. The SINOx control system uses a correlation function/map PEMS of the NOx mass flow, relative to the load, which was generated during engine bench tests. The reducing agent injection rate can be increased or decreased, depending on the required NOx yield and/or the measured NOx concentration downstream of the catalyst via a feedback control.

The SCR Process Aqueous urea is converted into ammonia through hydrolysis downstream of the injection nozzle. In the presence of the highly reactive, homogeneous, extruded SINOx (SCR) diesel catalysts, the nitrogen oxides (NOx) react with the ammonia (NH3) to yield nitrogen (N2) and water vapor (H2O). The operating temperature range is approximately 300–1020F. An oxidation catalyst is not required if an SCR diesel catalyst is used, because the fuel-efficient heavy-duty diesel engines practically produce no CO and very little VOC/HC and PM, which are further reduced by 80–90% and 20–30%, respectively. Oxidation catalysts are even undesirable if sulfur-containing fuel is used. The SINOx system is controlled automatically by a programmable microprocessor-based process logic control system. Active monitoring or supervision of the system by the operator is not required.

Summary on a Urea SCR Application in a Coal-Burning Utility Operation, American Electric Power Gavin Plant at Cheshire, Ohio Rated 2 X 1300 MW American Electric power (AEP), an international energy company based in Columbus, OH is capable of producing 38,000 MW electricity in 11 U.S. states and has a customer base of more than 4.8 million. AEP and other power generators in the Midwest and Southeast are required by court order to reduce their NOx emission by May 2003. AEP’s Gavin Plant has two 1300-MW generating units and is the largest utility station in Ohio This AEP Urea SCR project is designed to reduce NOx by approximately 70% to 0.15 lbs./pBTU of the coal-burning plant’s utility boilers at a cost of US$ 175 million. This project uses urea rather than anhydrous or aqueous ammonia as the reducing agent to overcome local safety concerns about the transport, handling, and storage of ammonia. (See Fig. 40.)

Acknowledgments The author would like to thank Siemens powergeneration’s KPW Group, Redwitz, Germany and Alpharetta, GA (USA) for supporting this project. The author is especially thankful to Dr. Juergen Zuerbig, head of Siemens’ SCR Business world-

148

Selective Catalytic Reduction wide and Dr. Raimund Mueller,* Head of the U.S. SCR Business and their staff for the support received and for the permission to incorporate Siemens material, used in prior papers and public presentations for this project. Without sharing their inside information about the European advancements in the SCR technology, in particular in mobile SINOx diesel catalyst applications in recent years, this project would not have been possible. Also, the author would like to thank EESI/Steuler for the support received and the permission to use material from public and internal documentation. The author is especially thankful to Nick Detor, Engineering and Project Manager of EES Inc., Cerritos, CA and Hans J. Wagner, Senior Manager of Steuler’s equipment Division, Hoehr–Grenzhausen, Germany who provided valuable information about some CER-NOx zeolite SCR catalyst applications. Furthermore, the author would also like to thank Dr. Ray Anthony, professor and head of the Chemical Engineering Department, Texas A&M University and his staff for their valuable guidance and editing effort. Finally, the author would like to thank the many professionals in the air pollution control community in the United States and Europe, with whom he has been associated over the years. The information shared privately, at conferences, and while serving with the author on committees in California, such as the Scientific Review Committee on Best Available Control Technology (BACT) at SCAQMD, on IC engines at Cal. ARB, and at joined projects at the University of California, was most helpful.

Bibliography

H. Lueders, R. Backes, G. Huethwohl, D.A. Ketcher, R.W. Horreck, R.G Hurley, and R.H. Hammerle, ‘‘An Urea Lean NOx Catalyst System for Light Duty Diesel Vehicles,’’ SAE 95-2493 (October 1995). R. Mueller, H. Roemich, and M. Joisten, ‘‘An Experience Report on Reducing Emissions of Criteria Pollutants of Stationary and Mobile Sources,’’ ICAC Forum ’98, Cutting NOx Emissions. (1998). G. Fraenkle, C. Havenith, and F. Chmela, ‘‘Test Cycle Development EURO3 for HD Diesel Engines,’’ CSAT, 1996. ‘‘UREA–SCR Test Achieves Big Cuts in NOx with Fuel Savings over Competing Technologies,’’ Diesel Fuel News (April 1999). J. Koeser, ‘‘SCR-De-NOx Katalysatoren,’’ Vulkan Verlag, Essen, Germany, 1992. N. Fritz, R. Mueller, J. Zuerbig, and W. Mathes, ‘‘On-Road Demonstration of NOx Emission Control for Diesel Trucks with SINOx Urea SCR System,’’ SAE Report 1999-01-0111 (1999). S. Eidloth, H. Roemich, M. Joisten, and A. Silini, ‘‘Exhaust Gas Aftertreatment Systems Onboard Seagoing Vessels,’’ Marine Conference and Exhibition, Brussels, 1999. G. Lepperhoff, G. Huetwohl, Q. Li, and F. Pischinger, ‘‘Untersuchung der NOx-Reduzierung im Abgas von Dieselmotoren durch SCR-Katalysatoren’’ [NOx Emission Reduction of

* Dr. Raimund Mueller became Manager of Mobile SINOx (SCR) applications worldwide in early 2000.


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Diesel Engine Exhaust Gas, a Comparative Analysis of European and U.S. (SCR) Catalysts], Report, Technical University RWTH–Aachen, Germany (1996). Vols. 6 & 7.‘‘Encyclopedia of Chemical Processing and Design,’’ Marcel Dekker, Inc., New York, 1978. Intermacom A.G., various in-house studies, 1990–1999. F.O. Witzel, ‘‘Massnahmen fuer einen wirtschaftlichen und sauberen Betrieb der SchiffsDiesel-Motoren’’ [Methods to Operate Ship Diesel Engines Efficiently with Low Emission], VDI Rep. 15 (198), (1998). BACT Guidelines, South Coast Air Quality Management District (SCAQMD) (1998). M. Joisten and R. Mueller, ‘‘Experience with the Reduction of NOx Emissions from Diesel Cogeneration Plants and Industrial Incineration Facilities,’’ Power Gen Europe, 1995. M.K. Khair, ‘‘Demonstration of Advanced Emission Control Technologies Enabling DieselPowered Heavy-Duty-Engines to Achieve Very Low Emission Levels,’’ MECA 1998. J. Kolar and H. Gleis, ‘‘NOx-Minderung in Rauchgasen,’’ VDI, (1987). R. Mueller and M. Grove: ‘‘Advanced SCR System Technology for the Simultaneous Reduction of NOx, VOC/Airtoxics and PM Emission,’’ California Air Pollution Control Officers Association (CAPCOA) Engineer Symposium, 1999. J. Peckham, ‘‘80% NOx Reduction with 8.5% Fuel Penalty, (NOx) Adsorber Test,’’ Diesel Fuel News (October 1999). ‘‘Exhaust Emission Reduction in Existing Ships,’’ Diesel & Gas Turbine Worldwide (April 1995). F. Fleischer, ‘‘NOx Reduction Technology Challenges Marine Diesel Builders,’’ Diesel & Gas Turbine Worldwide. R.H. Thring and R. Hull, ‘‘NOx Control Technology Data Base for Gas-fueled Prime Movers, Southwest Research Institute, 1987. L. Bruce, ‘‘Reducing NOx Emission,’’ Power (1981). J. Davis and G. Duponteil, ‘‘Using SCR for NOx Control, Power (1986). J.H. Wasser and R.B. Perry. ‘‘Diesel Engine NOx Control with SCR,’’ EPA Stationary Source Symposium, 1984. C. Havenith and R.P. Verbeek, ‘‘Transient Performance of a UREA DeNOx Catalyst for Low Emission Heavy-Duty Diesel Engines,’’ SAE Paper 970185, 1997. Diesel Net, various reports 1998–1999. http:/ /www.dieselnet.com. Diesel Fuel News, various reports (1998/1999). N. Kato, N. Kokune, B Lemire and T. Walde, ‘‘Long Term Stable NOx Sensor with Integrated In-Connector Control Electronics,’’ SAE Paper 1999-01-0202 (1999). D. Simbeck, ‘‘The Future of Distributed Power Generation,’’ Power Generation in the 21st Century, Energy Frontiers & U.S. Department of Energy, (1997). W.R. Miller, J.T. Klein, R. Mueller, W. Doelling, and J. Zuerbig, ‘‘The Development of UREA-SCR Technology for U.S. Heavy-Duty Trucks,’’ SAE Paper 2000–01–0190 M. Kirchner, ‘‘Die Bedeutung von Ammoniak and Ammonium fuer Mensch und Umwelt,’’ GSF-Forschungszentrum, 1999. ‘‘Use of Ammonia in NOx Pollution Control,’’ ICAC, 1999. M. Grove and W. Sturm, ‘‘NOx Reduction with the CER-NOx (SCR) Process,’’ ASME Proceedings, South West Research Institute, October 1988. M. Grove and W. Sturm, ‘‘NOx Abatement System for Glass Melting Furnaces,’’ Ohio State University, November 1988. MANFRED GROVE

150

Nanophase Materials in Chemical Process


Introduction The selection of construction materials for chemical process equipment is a critical factor in the efficient operation of a chemical process plant. Engineering materials must satisfy a variety of criteria in order to operate successfully in a process environment. The mechanical properties of a material must be satisfactory with regard to tensile strength, elastic modulus, and resistance to fracture, wear, fatigue, and creep across a broad range of temperatures and/or pressures. The chemical property of corrosion resistance is also a significant factor to consider in process material selection. Additionally, the effects of process thermal conditions and temperature cycling on material properties are relevant to the durability of the material in a working plant. Beyond these generic material selection considerations, process-specific material properties are regularly required. The demand for new materials with an ever-expanding range of properties has led to the exploration of nanophase materials. Nanophase materials are in a unique portion of the ordered length scale of materials between molecular and bulk domains. As the domain size of a given material increases across the nanometer size regime of between 1- and 100-nm molecular properties become increasing collective and bulklike in nature. Thus, many bulk properties of a material considered to be inherent constants are, in fact, size-dependent variables across the nanophase size regime. The practical effect of material properties being adjustable across the nanophase size regime is that material domain size becomes another mode of modifying process material properties, along with the conventional modes of composition and treatment conditions. This article highlights nanophase material developments relevant to chemical process construction and operation. Materials science has experienced substantial progress in the synthesis and characterization of nanophase materials [1]. Nanophase materials have properties which vary from those of the bulk including reactivity [2], magnetism [3], melting temperature [4], and mechanical strength [5]. Although the high surface area of nanophase materials has long been exploited in catalysis, promising results suggest broad application of nanophase materials in all aspects of future chemical processes. Nanophase concepts and technologies can be applied in many different configurations, including active nanoclusters, nanocrystalline materials, nanocomposites, nanotubules, and even ‘‘nanoparticles of vacuum,’’ in which desired materials properties are obtained by introducing nanometer-scaled voids or pores in materials [6,7]. These structural configurations combine with the fundamental aspects of nanophase materials to provide tools for the tailoring of desirable mechanical and chemical properties.


151

Nanophase Properties: Origin and Overview Nanophase properties are largely a function of an incomplete electronic band structure and surface sites comprising a non-negligible percentage of the total atoms. These characteristics of nanophase systems will be manifest and helpful in understanding the material properties discussed in this article.

Molecular Origin The central effect of an incomplete electronic band structure in a system is that the smaller the dimensions, the higher the energy of the lowest excited state. From a philosophical standpoint, nanocrystals can be examined as either a summation of a large, but finite, number of interacting atoms or as a small piece of material excised from a bulk crystal. The former approach relies on atomic orbital hybridization to build a band structure. Because the evolution from atomic orbitals to the band structure associated with the bulk follows directly from atomic hybridization, a linear combination of atomic orbital–molecular orbitals (LCAO–MO) method is invoked to explain nanophase behavior. Changes in electronic structure as a function of size are most easily illustrated by the use of a one-dimensional system [8]. An infinite chain of N ‘‘atoms,’’ each possessing one π orbital, containing one electron and separated by a distance, a, in the absence of any interatomic interactions, has N degenerate orbitals, φ i, each with an orbital energy, e, given by e ⫽ ⬍ φ i |H |φ i ⬎

(1)

where H is the electronic Hamiltonian. When the individual units interact, the resonance energy is given by β ⫽ ⬍ φ i | H|φ j ⬎

(2)

The result of a unit of the chain interacting with adjacent units is to lift the degeneracy of orbitals, as shown in Fig. 1. The energy levels of these one-electron orbitals are derived using the Hu¨ckel approximation of neglecting orbital overlap: E( j, N)⫽ e ⫹ 2β cos

冢

冣

jπ , j ⫽ 1, 2, . . . , N (N ⫹ 1)

(3)

⫽ e ⫹ 2β cos(k j a) where k j ⫽ j π/(N ⫹ 1) a. As N approaches infinity, k, the wave vector, and E(k) become continuous variables from 0 to π/a and e ⫺ 2β to e ⫹ 2β, respectively.

152


FIG. 1 The result of adjacent orbital interactions serves to lift the degeneracy of the orbitals resulting in the energy levels of N degenerate orbitals splitting into a band.

It is now apparent that π/a is the limit of the Brillouin zone and each ‘‘atom’’ in the chain is comparable to a unit cell in the semiconductor. For a finite chain, the eigenvalues are discrete and energetically separated by increasing amounts as the chain shortens, as shown in Fig. 2. The extreme eigenvalues asymptotically approach e ⫾ 2β as the chain length increases. This model system shows how adjacent eigenvalue spacings increase as the system shrinks, as in the case of a metal, such as Pt [9]. The addition of a twounit repetition pair and Jahn–Teller distortion serves to open a gap between the highest occupied and lowest unoccupied molecular orbitals (HOMO–LUMO) in the middle of the dispersion curve at k ⫽ π/a′, where a′ is the repetition pair spacing, as shown in Fig. 3. The dependence of the electronic properties on particle size shows dramatic changes between the dimensions of atoms and the bulk material [9]. As the preceding model suggests, the electronic band structure incrementally grows toward the

FIG. 2 The band structure for a one-dimensional chain of N atoms, each with a single electron and an interatomic separation a. Energy in units of overlap energy. (—): an infinite energy chain; 䊉 N ⫽ 9 eigenvalues; ■ N ⫽ 5 eigenvalues.


153

FIG. 3 The effect on a uniform chain of the introduction of a two-atom repeat unit and a Jahn–Teller distortion to produce a HOMO–LUMO gap at k ⫽ π/a′.

bulk as the system size increases. Likewise, one might infer that bulk thermodynamic characteristics arise from an additive stabilization of atomic properties.

Surface to Volume Ratio The surface of a nanocrystal contains a non-negligible fraction of the total atoms. For example, a 1-nm-radius nanocrystal of arbitrary composition contains about 200 atoms, nearly 60% of which occupy surface sites. The effect of a surface can be incorporated into thermodynamic models to account for the modified behavior associated with fine particulate metals and molecular clusters. The surface free energy produces a size dependence in the chemical potential, µ, of the particle. For a constant density liquid droplet, the chemical potential is given by µ ⫽ µ∞ ⫹

2γ ρR

(4)

where γ is the surface tension, ρ is the density, and R is the droplet radius. The same equation describes crystalline solids with shapes satisfying the Curie–Wulff equation: γi ⫽ construction. hi

(5)

where γ i is the surface energy of face i and h i is the distance from that face to the particle center of mass.

154

Nanophase Materials in Chemical Process Chemical bond energy and coordination differences also exist between surface and bulk lattice sites within a material [10,10a]. With the large percentage of constituent atoms or molecules of a nanophase material present on surface sites or within a few atomic layers of the surface, nanophase materials often are understood as a manifestation of surface properties which are masked in studies of bulk materials. By way of example, Cu nanoparticles show a significant change in interatomic spacing, owing to surface tension effects and strained surface geometries associated with nanophase systems, with the spacing increasing from 2.22 to 2.50 to 2.56 between Cu 2, 0.5-nm nanoparticles, and the bulk, respectively [11]. Properties

Melting Melting is a fundamental, but not simple, process of a crystalline solid at a distinct temperature, T m undergoing a phase transformation to a liquid. The complexity in this seemingly mundane process lies in the atomistic nature of the transformation. Although thermodynamics rigorously defines T m as the temperature at which the Gibbs free energy of the solid and liquid phases is equivalent, there is no specification as to the mechanism of melting. Thermodynamics is also mute as to the kinetic issue of the rate at which the transformation occurs. As such, thermodynamic theories of melting assume a homogenous process occurring simultaneously throughout the system. Because all real solids contain extrinsic (thermodynamically metastable) defects such as surfaces, dislocations, and grain boundaries, which often are both thermodynamically and kinetically more reactive to transformation, it stands to reason that they play a significant role in melting. The loss of long-range order is a characteristic of melting, and in the high concentration limit, extrinsic defects represent a similar disruption of the long-range order. Metallic and semiconductor nanocrystals have been shown to possess reduced melting temperatures when compared with the bulk. The depression in melting and annealing temperature is evident throughout the nanocrystal size regime, but most profound in the size range from 1 to 6 nm. Melting studies on a range of nanocrystals have established that the melting temperature is size dependent in the nanometer-size regime and is approximately proportional to the inverse particle radius, regardless of the material identity, as shown in Fig. 4. The size-dependent melting temperature of metallic nanocrystals has included studies of Au [12], Pb and In [13], Al [14], and Sn [15]. Even though semiconductors have directional and somewhat covalent bonds and the crystal facets of binary semiconductors do not all exhibit bulk stoichiometry, the size-dependent melting temperature has been shown in the direct band-gap semiconductors CdS [16] and GaAs [17]. Nanocrystals of the indirect semiconductor Si also exhibit size-dependent melting [18]. The reduction in melting temperature as a function of nanocrystal size can be enormous. For example, 2-nm Au nanocrystals melt at about 300°C, as compared to 1065°C for bulk gold [12].

Softening of Nanocrystalline Metals The properties of bulk materials fail to hold true when dealing with nanophase materials. For example, nanocrystalline solids often are sought for their increased


155

FIG. 4 The melting temperature of materials as a function of particle size, plotted as a fraction of nanocrystal melting temperature T m to the bulk melting temperature T b.

hardness and/or ductility [19]. The Hall–Petch effect is observed in polycrystalline metals with random grain orientations, where both hardness and yield stress typically increase as the grain size decreases. Yet contrary to the Hall–Petch effect, computer simulations of nanocrystalline copper show that as the grain size decreases down to values less than approximately 7 nm, hardness and yield decrease as the grain size decreases [19].

Nanophase Materials Intermetallic Compounds Nanophase materials technology can be used to reinforce an intermetallic compound with a ceramic to obtain desirable characteristics of each type of material. For example, the reinforcement of titanium and aluminum intermetallic compounds (Ti 3Al, TiAl) with TiC improves the toughness and retention of strength across a broad range of temperatures while continuing to provide a lightweight material with good oxidation resistance [20]. Mechanical alloying is a viable method for homogeneously distributing nanophase particles of TiC in an amorphous Ti–Al matrix [20], thereby creating one such in situ ceramic-reinforced intermetallic compound. Ceramic reinforcement SiC–Si 3N 4 composite powders have been synthesized through a sintering process, resulting in nanophase SiC particles distributed intergranularly and intragranularly throughout the Si 3N 4 matrix [21]. This resulted in a dramatic improvement in high-temperature strength and creep resistance [21].

156


Chemiresistors High-performance chemiresistors have been developed through the use of a colloidal nanophase structure known as a metal–insulator–metal ensemble (MIME) [22]. The resistance of a chemiresistor changes when the device is placed in the presence of specific chemicals, as both adsorbed and absorbed vapors affect the transfer of electrons across the matrix [22]. A thin transducer film having gold nanoclusters terminated with alkanethiol ligands is deposited onto an interdigital microelectrode [22]. By selecting the absolute and relative sizes of the gold nanocluster and the ligand thickness, chemiresistors can be developed to target-specific chemical species while ‘‘ignoring’’ others. Thus, one can selectively detect the presence of various chemicals through changes in electrical resistance.

FIG. 5 Vapor response isotherms of the Au: C 8 (1 : 1) MIME sensor to toluene, tetrachloroethylene (TCE), 1-propanol, and water, based on 15°C vapor pressures. The inset displays the toluene response down to a 2.7 ppmv vapor concentration. (From Ref. 22.)


157

FIG. 6 A simple two-dimensional sketch of the gold cluster film morphology illustrating the gold core, alkanethiol ligand shell, and region of lower ligand density. (From Ref. 22.)

One such construct has been shown to be extremely sensitive to toluene and tetrachloroethylene (TCE) vapors, but insensitive to water vapor [22]. Such a chemiresistor serves as the basis for a new class of miniature chemical sensing devices that remain unaffected by typical environmental conditions, such as humidity [22].

Polymer Films Although the coating of lenses is an old practice, high-performance broad-band antireflection coatings have not been commonplace due to materials limitations [23]. Antireflective (AR) coatings function through destructive interference of reflected light (see Fig. 7A) [23]. The mathematics of this destructive interference, combined with a review of the available dielectrics, show that single-layer antireflective coatings cannot possess the desired refractive indices [23]. However, nanoporous films, with pore sizes much less than the incident light wavelength can be tailored to specific refractive indices. The refractive index is dependent on the pore volume ratio of the film [23]. Nanoporous films have been prepared by demixing a binary polymer blend during spin-coating, and then using a selective solvent to remove one of the source polymers [23]. Nanocluster technology is being applied to embed metal particles in polymer

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FIG. 7 (A) Reflection of light from both interfaces of an AR layer. For a given wavelength and incidence angle, light transmission is maximized when the two reflected beams interfere destructively. (B) Preparation of a binary polymer film. Initially, both polymers (black and gray) and the solvent form one phase. During spin-coating, phase separation sets in, and after evaporation of the solvent, a lateral phase morphology is obtained. (C) The film is exposed to a solvent that is selective for one of the polymers, producing a porous film. (From Ref. 23.)

films in order to provide new properties of strength, magnetism, or infrared absorption capabilities [24]. Iron, iron oxide, silver, and copper have been placed successfully into permanent polymer lattice sites [24]. In such constructs, a small quantity of embedded metal allows the film to remain transparent while still affecting its performance characteristics [24]. Applications for such materials include magnetic ‘‘watermarks’’ and heat-absorbent window coatings [24].

FIG. 8 Atomic force microscope (AFM) images of two porous PMMA films ⬃110 nm thick. After spin-casting of a PS–PMMA–THF mixture onto silicon oxide surfaces, the PS phase was removed by washing the sample in cyclohexane. (A) Films prepared from higher M w PS and PMMA show average structure sizes of ⬃1 µm. (B) If low M w PS and PMMA are used, the lateral structure size is reduced to ⬃100 nm. Although the film in (A) appears opaque, the nanoporous film in (B) is transparent, with a low effective refractive index. (From Ref. 23.)

Nanophase Materials in Chemical Process TABLE 1

159

Summary of the Magnetic Data Obtained on Four Different Samples

Sample Co Co: Pt Co: Pt3 Au at Co: Pt

Tb (K)

Hc (G)

µ(eff) (emu/g)

70 140 130 30

1800 2700 2000 1000

0.115 0.0598 0.0357 0.0166

Magnetics As the information revolution continues, the need for reliable, more compact storage media continues to grow. As researchers look to alloys to improve upon the magnetic properties of traditional materials, nanophase metallic alloys are viewed as attractive candidates because of its ability to control properties as a function of domain size [25]. The magnetic properties of nanophase cobalt, cobalt–platinum, and gold-coated cobalt–platinum appear to be promising high-performance materials showing dramatic effects in the blocking temperatures (T b ) and coercivities (H c ) (see Table 1) [25]. High-density media storage requires the nanoparticles to approach monodispersity and a defect-free state. Nanometer-scaled ferromagnetic particle arrays have reversible rotation and magnetic switching behaviors associated with magnetic core diameters and shape anisotropy that affords a variety of new device design options [26]. At temperatures above 30 K, magnetization reversal is dependent on thermal activation within a volume that increases with particle diameter [26]. Such a model is useful in predicting blocking temperatures [26].

Catalysts

Properties The ability to conduct molecular studies of surface behaviors of nanometer-sized particles and/or nanoporous materials has led to an understanding of surface effects in catalysis [27]. Improved scanning tunneling microscopy (STM), sum frequency generation (SFG), vibrational spectroscopy, and atomic force microscopy (AFM) have provided information about surface reactions and catalysis at relatively high pressures, as compared to previous instrumentation [27]. The ability to produce nanometer-scaled ordered arrays and then to conduct detailed studies of their surface behaviors at, or near atmospheric pressures approaching actual catalysis conditions has led to a greater understanding of the pressure dependence of metal nanocluster catalysis [27]. SFG studies at high pressures reveal that CO molecules occupy binding states on Pt(111) that are not present at low pressures, whereas adsorbates that are readily detectable in low-pressure surface studies often are stationary spectators or totally absent during high-pressure catalytic reactions. Further, turnover rates are negligible compared with the overall reaction rate [27]. These findings illustrate the need for pressure-specific studies of the catalytic be-

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Nanophase Materials in Chemical Process havior of metal nanoclusters, because the reaction pressures greatly influence the outcomes in a way that cannot be extrapolated from low-pressure studies [27]. In additional, the level of surface defects, or other irregularities, can have similarly profound effects on catalysis behaviors [27]. At the same time, investigations continue into the ‘‘living metal polymer’’ model, in which nanocluster growth is considered analogous to polymer growth [28] and autocatalytic surface growth once M(0) nanoclusters have been nucleated [29]. Of interest are the developments in size-control and size-prediction methodologies, which have been tested successfully in monometallic nanoclusters such as iridium(0) [29]. Such models are predicted to be applicable to bimetallic, trimetallic, and multimetallic transition metal nanocluster systems [29]. An example of this process of slow, continuous nucleation, followed by autocatalytic surface growth, is shown in Scheme 1 and Fig. 9. It is the ratio of the rate of growth to the rate of nucleation [R ⫽ k 2 (nanocluster active sites)/k] [29] that can be used to predict, and consequently control, nanocluster size [29]. The research into these methodologies also has led to an endorsement of the ‘‘magic

(a)

Nucleation k1

A → B (1) (2) (3)

(COD)Ir(P 2W 15Nb 3O 62 )8⫺ ⫹ 2 acetone (COD)Ir(acetone) 2⫹ ⫹ P 2W 15Nb 3O 629⫺ (COD)Ir(acetone) 2⫹ ⫹ 2.5H 2 → Ir(0) ⫹ H⫹ ⫹ S ⫹ 2 acetone nIr(0) → Ir(0) n (b) Autocatalytic Surface Growth k2

A ⫹ B → 2B (4)

(COD)Ir(P 2W 15Nb 3O 62 )8⫺ ⫹ 2.5H 2 ⫹ Ir(0) n → Ir(0) n⫹1 ⫹ H⫹ ⫹ P 2W 15Nb 3O 629⫺ ⫹ S

Net Reaction: 300 [(COD)IR(P 2W 15Nb 3O 62 )8⫺] ⫹ 750H2 → Ir(0) 300 ⫹ 300 P 2W 15Nb 3O 629⫺ ⫹ 150 H⫹ 300S

SCHEME 1 Minimum mechanism for the formation of Ir(0) nanoclusters, consisting of (a) slow, continuous nucleation (steps 1–3), rate constant k 1 for the pseudoelementary step, A → B, followed by (b) fast autocatalytic surface growth (step 4), rate constant k 2 for the pseudoelementary step A ⫹ B → 2B. Nucleation and growth are separated in time because k 1 ⬍⬍ k 2 [B], which, in turn, is a key to the observed formation of a nearmonodisperse (⫾ ⱕ15%) particle size distribution. (From Ref. 9.)


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FIG. 9 Idealized, roughly-to-scale representation of a P 2W 15Nb 3O 629⫺ polyoxoanion and Bu 4N⫹ stabilized Ir(0) ⬃300 nanocluster, [Ir(0)⬃300(P4W30Nb6O12316⫺)⬃33](Bu4N)⬃300Na⬃228. The Ir(0) atoms are known (by electron diffraction) to be cubic close packed as shown. For the sake of clarity, only 17 polyoxoanions are shown, in their monomeric form, and the ⬃300 Bu 4N⫹ and ⬃228 Na⫹ cations have been deliberately omitted. (From Ref. 29.)

number’’ theory of nanoclusters, which states that because closed-shell structures are stable, they will be more common in the size distribution of completed nanocluster formations [29]. Further exploration of the catalysis properties of transition metal colloids indicates that their behavior is extremely reaction-specific and dependent on many factors. For example, in the synthesis of silicone polymers, the catalytic performance of bimetallic colloids of Au(core)/Pt(ligand) and Pd(core)/Pt(ligand) was compared to the performance of Pt nanoclusters formed during the induction period of typical industrial reactions [30]. Previously, Au–Pt had been shown to provide marked improvement over Pt alone in the semihydrogenation of 2-hexyne into cis2-hexene [31]. However, in the synthesis of bis(trimethylsiloxyl)octamethylsilane (BTMOS), the Au–Pt showed no improvement over Pt alone, whereas Pd–Pt showed marked improvement [30]. These experiments were designed in order to study the behavior of a bimetallic colloid with a more electronegative core than the ligand (Au–Pt) and compare it to that of one with a more electropositive core than ligand (Pd–Pt) [30]. For this particular commercial synthesis, the more electropositive core seemed to yield superior results. However, in light of the Au–Pt semihydrogenation mentioned, experimental verification is required before such findings can be extended to other catalytic systems [30]. Nanocluster agglomeration is prevented through surface stabilization using ligands, polymers, or surfactants [32]. Using a combination of STM and transmis-

162


FIG. 10 A schematic diagram showing how STM and TEM can be used in concert to determine surfactant stabilizer thickness. Thicknesses observed in this manner are consistent with standard MM2 force-field calculation theoretical values. (From Ref. 32.)

sion electron microscopy (TEM) to observe and examine the size relationships between surfactant stabilizers and metal cores, it has been shown that shell thickness is independent of core size and is directly dependent on the size of the surfactant ion, as shown in Fig. 10 [32].

Synthesis Any discussion of nanophase materials in chemical processes must address the topic of catalysis. The high surface area to volume ratios associated with nanophase materials makes them exceptional catalysts as compared to bulk materials. Although nanophase metal catalysts supported on substrata have been industry standards for high-throughput reactions, in recent years particle size control has strived to produce stoichiometrically defined particles. Such a well-defined nanophase catalyst offers the prospect of engineering the ratio of edge atomic surface sites on a particle surface to alter reaction product efficiencies and mixtures. To this end, transition metal nanocluster systems have been synthesized which behave as isolable and compositionally well-defined soluble heterogeneous catalysts [33]. Further, new methodologies and a better understanding of nanoclusters in general are calling into question previously identified catalysts in previously explored reactions [33]. One such example is a hydrogenation of benzene, in which the accepted ionpair catalyst, [(C 8H 17 ) 3NCH 3]⫹[RhCl 4]⫺, has been discredited, with the catalyst shown more likely to be Cl⫺- and [(C 8H 17 ) 3NCH 3]⫹-stabilized Rh(0) nanoclusters [33]. This indicates that in the cases of other putative homogenous catalysts where a facile heterogeneous M(0) catalyst is well established, catalysis by even trace amounts of possibly highly active nanocluster catalysts cannot be ruled out [33]. The development of industrial catalysts,many of which are composed of metal particles on oxide substrata, is a topic of great commercial interest [34]. Fabrication and engineered control in order to deposit tailored metal clusters on oxide surface with an ordered structure remain a goal of catalyst research [34]. Lithographic


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nanostructure fabrication technologies of the semiconductor industry can be applied to create such engineered clusters and to provide viable model systems for industrial-supported catalysts [34]. This model of catalytic structure and behavior has been tested against ethylene hydrogenation on a platinum nanocluster [34]. In a study of the thermal stability of supported silver catalysts, these model systems and fabrication techniques were used to demonstrate that the presence of oxygen is the key factor influencing the thermal stability of the silver nanoclusters [34]. No migration of silver clusters was observed at ⬎700°C in the absence of oxygen [34]. Yet, with increasing annealing temperature in oxygen, silver cluster surface oxidation occurs at ⬍200°C [34]. Methods of nanostructure lithographic fabrication also can be applied to other surface science studies. When combined with AFM, these developments are used to determine surface mechanical properties, such as the elastic modulus, on nanometer-scaled samples [34]. Methodologies also have been developed to measure the yield of an ion-sputtering process [34]. In addition to the synthesis and use of ligand-stabilized metal clusters, significant benefits are realized by polymer stabilization of nanocluster catalysts [35]. The platinum nanoclusters’ catalytic performance on polymer-based supports has been compared to performance on oxide supports [35]. Polymeric supports show a marked increase in the stability of the catalysts, especially at room temperatures [35]. Although stabilizers control particle size and prevent agglomeration, cluster surface passivation is often the result, with the net effect of reducing catalytic performance [36]. One proposed solution to this dilemma is the use of dendrimers to act simultaneously as monodisperse synthesis templates and stabilizers [36]. Researchers have partitioned transition metal ions, such as Pt, into the interior of polyamidoamine (PAMAM) and have achieved an encouraging combination of particle size control, particle stability, and electrocatalytic behavior and performance [36]. (See Fig. 11.) Nanophase catalysts also function in decomposition reactions of harmful greenhouse gases,such as CO 2[37]. Experiments with different ferrites (XFe 2O 4 ) have shown that with appropriate selection of X and particle size, CO 2 is broken down into carbon and oxygen with virtually no CO by-product [37]. The decomposition of CO 2 also produces quantities of methane. Comparisons of Zn, Ni, and Co ferrites and of varying particle sizes yielded dramatic results, with ZnFe 2O 4 showing the most promise [37]. In addition to purely chemical catalysis, nanocluster materials also function as electrocatalysts. In the synthesis of sodium chlorate, the Dimensionally Stable Anode (DSA) has realized extensive energy savings on one side of the oxidation reaction, but high activation overpotentials on the cathodic side still contribute to large energy losses [38]. The experimental production of a solid Ti 2RuFe nanocrystalline cathode has shown promising results. In fact, once the nanocrystalline material was pressed into a usable cathode, the activation overpotential was reduced dramatically when compared to a standard iron electrode [38]. However, this method of electrode fabrication failed to produce anticipated current density increases expected from the very small crystal size [38]. Template methods, combined with chemical vapor deposition (CVD), also can be used in the preparation of carbon nanotubules, with diameters ranging from 20

164


FIG. 11

Schematic illustration of fourth-generation (G4) PAMAM dendrimers having EOH and NH 2 terminal groups, synthesis of Pt nanoparticles within the hydroxyl-terminated dendrimer template, and attachment of the composite to an electrode surface. Between 12 and 60 Pt2⫹ ions can be loaded into a single dendrimer and, upon reduction with BH 4⫺, an entrapped cluster containing the same number of atoms results. The dendrimer-encapsulated Pt nanoparticles are electrocatalytically active. (From Ref. 36.)

to 200 nm [39]. Such tubules can be filled with metal catalyst nanoclusters of the types previously discussed to display interesting and potentially useful electrocatalytic properties [39]. A basic method used to synthesize metal-nanocluster-filled carbon nanotubes includes carbon deposition by CVD onto an alumina template membrane, followed by immersion in a metal ion solution. Air-drying and reduction by hydrogen gas completes the formation of the metal nanoclusters within the nanotubules. Alumina is removed by HF immersion [39]. The mechanical and electrochemical properties displayed by metal-filled nanotubules hold promise for the fuel cell industry [39].

Glass Silicates A physicochemical analysis of nanophase crystalline silicalite-1 shows that it has many properties in common with micrometer-sized crystalline silicalite-1 [40]. These common properties include a refined structure and concentrations of tetrapropylammonium species incorporated during the synthesis [40]. Yet, nanophase silicates have nondegenerate spectroscopic features such as the characteristic framework infrared (IR) vibration (550 cm⫺1 ) of the micrometer-sized crystal splitting into a doublet (at 555 cm⫺1 and 570 cm⫺1 ) in the nanophase material [40]. In addition, the nanophase material exhibits a high concentration of defect sites, a


FIG. 12

165

Log-log plot of the bulk (K), shear (G), and Young’s (E) moduli for bulk and nanophase a-SiO 2, as a function of the density relative to the bulk density. The solid lines are the best least-squares fits for each of the moduli. (From Ref. 41.)

strain in the crystallites along the a crystallographic direction, as well as a twostage dinitrogen physisorption in the low-pressure region [40]. The pore morphologies and mechanical behaviors of nanophase amorphous SiO 2, investigated by molecular-dynamics (MD) simulations, show that the bulk amorphous densities of the various nanophase a-SiO 2 glasses are characterized by different pore sizes and distributions, yet the morphology of the pores, defined in terms of the fractal dimension of pores and the roughness exponents of pore–silica interfaces, is similar across various densities [41]. This consistent short-range order (SRO) of nanophase silica glass of various densities differs little from the SRO of bulk glass, with both structures consisting of corner-sharing Si(O 1/2 ) 4 tetrahedra [41]. However, analysis of the first sharp diffraction peak (FSDP) shows a significant difference in the intermediate-range order (IRO), dependent on the density, which is controlled synthetically in the nanophase silica glass [41]. In terms of mechanical properties, the elastic moduli of both the bulk and the nanophase a-SiO 2 clearly are density dependent [41], as shown in Fig. 12. The moduli (M) scale as M ⬀ (ρ-bar)3.5 ⫾ 0.2, where ρ-bar is the ratio of the sample density to the density of bulk silica glass [41]. An understanding of this powerlaw dependence can lead to very specific tailoring of physical and mechanical properties for various uses.

Glass Ceramics Nanophase glass–ceramic technology also is a rapidly growing field with a wide variety of commercial applications. Nanophase microstructures composed of crystals ⬍100 nm in size are achieved through efficient nucleation and slow crystal growth, providing an impressive uniformity of microstructure and the controlled mechanical properties not available in glass–ceramics of larger crystalline microstructures [42]. Transparent nanophase glass–ceramics achieve near-zero coefficients of thermal expansion, along with high thermal stability, high thermal shock resistance, and, of course, transparency [42]. The requirements to achieve transparency have previously been recited [42]. Glass–ceramics with these properties typically utilize lithium-stuffed β-quartz crystals to satisfy scientific and commercial applications,

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Nanophase Materials in Chemical Process such as telescope mirror blanks, stove cook tops, cookware, woodstove windows, fire doors, and other technical devices [42]. Near-zero coefficients of thermal expansion result from adjusting the various composition components and percentages in stuffed lattices [42]. ‘‘Stuffed’’ βquartz crystals are so named because Al3⫹ replaces Si4⫹ in the β-quartz framework of interlinked helixes of SiO 2 tetrahedra, with the charge balance being maintained by ions that stuff the interstitial tetrahedral cavities [42]. Other interesting subclasses of transparent glass–ceramics, many of which also have a ‘‘stuffed’’ microstructure, consist of materials with coefficients of thermal expansion near that of silicon [42], materials such as transparent mullite or spinel glass–ceramics that can serve as superior-performance host media for luminescent transition metal ions such as CR3⫹ [42], and oxyfluoride glass-ceramics used as hosts for optically active rare earth (RE) cations, because of their low phonon energies and broad transparency in the IR region of the spectrum, used for the amplification of light in telecommunications systems [42].

Conclusions As we have seen, nanophase materials show great promise for many areas of chemical process. Their unique properties, brought about by a high surface area to volume ratio and incomplete electronic band structures, give them extraordinary flexibility with regard to the desired tailoring of mechanical and chemical properties. Composition and configuration options offer exciting opportunities to take nanophase materials out of the laboratory and into innumerable industrial applications.

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Process Safety and Risk Management Regulations: Impact on Process Industry

Introduction Chemical accidents have been caused by a number of reasons, including human error, design flaws, lack of process and engineering knowledge, equipment failure, and natural disasters. The danger posed to the employees of a chemical plant as well as the public is illustrated by the accidents that have occurred in onshore as well as offshore chemical process industries. Figure 1 provides accident statistics for 1989 from the Accidental Release Information Program (ARIP) of the U.S. Environmental Protection Agency [1]. ARIP statistics cover catastrophic and unplanned releases of chemicals into the atmosphere. However, these statistics underline the fact that a large number of accidents and catastrophic releases occur because of design flaws, wrong equipment specifications, and lack of or disregard of operating and maintenance procedures. The boardroom perspective on the cause of these accidents and what to do about them varies. The total number of process plant accidents cannot be estimated accurately because of underreporting. However, it is clear that the number of accidents is large and many people, both workers and the public, are affected adversely by these accidents. For example, in 1991, the National Response Center received over 16,300 calls reporting the release or potential release of hazardous chemicals [3]. Another study [4] analyzed the EPA’s Emergency Response Notification System database of chemical accident notifications. The study found that from 1988 through 1992, an average of 19 accidents

Process Safety/Risk Management Regulations

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FIG. 1 U.S. Environmental Protection Agency Statistics Accidental Release Information Program— 1989. (Reprinted with permission from Ref. 2.)

occurred each day (i.e., more than 34,500 accidents involving toxic chemicals occurred over the 5-year period). The promulgation of the Toxic Release Inventory Reporting requirements [5] as part of the Clean Air Act Amendments of 1990 led to the submittal of toxic release information which clearly delineated the number and extent of toxic chemical releases and their potential impact on the public and the environment. In addition to the industry and government agencies, the university has a critical role in changing this situation. In addition to statistics and the sheer number of facilities involved, a number of highly publicized chemical plant incidents in the 1970s and 1980s focused attention on management systems and technologies. For example, the 1974 Flixborough accident occurred because a temporary pipe was used to replace a reactor which had been removed for repair [6]. The temporary piping was not properly designed and supported merely on scaffolding. A management of change system could very well have prevented the incident. The causes behind the 1984 Bhopal accident, which involved the release of methyl isocyanate and caused thousands of fatalities, have been investigated quite extensively with varying conclusions. However, the need for inherently safer design considerations is quite unanimous. Bhopal and many other process plant incidents including the 1984 Mexico City disaster [7] also emphasize the need for application of structured management systems for hazard recognition and identification. According to the official report [8] following the 1988 Piper Alpha disaster in a North Sea offshore rig, a pump relief valve was removed for overhaul and the open end blanked. Another shift not knowing the relief valve was missing, started up the pump. However, this primary reason does not obviate the fact that a number of other factors contributed to the extensive damage. Among other things, the Phillips 1989 explosion [9] in the high-density polyethylene plant demonstrates the need to adhere to operating procedures and implementing appropriate management systems for contract workers. Many process plant accidents in the 1970s and 1980s also exposed the need for management systems to ensure process and equipment integrity. Change in population demographics, increasing awareness of process plant

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Process Safety/Risk Management Regulations hazards, and, above all, the continuing threat of a chemical catastrophe continue to provide the impetus for governments to develop legislation for eliminating or minimizing the potential of such accidents. International efforts include the Seveso Directive covering members of the European Community. Many other nations also have similar laws, such as the Sedesol guidelines in Mexico for performing process risk audits and the post-Bhopal accident prevention law in India. The World Bank has developed guidelines for identifying and controlling hazards, and the International Labor Organization has developed a code of practice for preventing major accidents. In 1990, the U.S. Congress enacted the Clean Air Act Amendments (CAAA) into law. The CAAA directed the Occupational Safety and Health Administration (OSHA) and the Environmental Protection Agency (EPA) to develop regulations to reduce the frequency and severity of chemical plant accidents. In keeping with the congressional mandate given in Section 304 of the CAAA, OSHA promulgated the Process Safety Management (PSM) regulation on February 24, 1992. The PSM regulation is intended to protect workplace employees. Similarly, as mandated by Section 301(r) of the CAAA, EPA promulgated its risk management program regulation in 1996, to protect the public and the environment. In the United States, federal agencies are not the only government regulators active in the chemical accident prevention arena. Several states have empowered their health, safety, and environmental agencies to create regulations requiring companies to establish and practice specific programs to improve safety.

The Impact of the Industrial Revolution The eighteenth century was the beginning of technological development, which affected society and commerce in ways that are felt even today. This technological development, known as the industrial revolution, was one of the main revolutions of this era. The industrial revolution changed manufacturing by changing the way people worked. For one thing, it brought work out of the home and centralized it in the beginning to small and simple plant operations and increasingly to large and complex operations. The industrial revolution grew more powerful each year as new inventions and manufacturing processes added to the efficiency of machines and increased productivity. The industrial revolution has had far more impact on the world than the political revolutions of the era, because the Industrial Revolution effects on society are longer lasting. For example, today we have automobiles, television, and computers, all made possible by this revolution. Without the industrial revolution, we would not have the technologies that we have today and neither would we have the standard of living that we enjoy. Before the introduction of machines and the factory setting, goods were manufactured by hand, in single homes or cottages, where the owner worked side by side with his employees. This changed with the introduction of machines and mass production. The major advances in technology, particularly in the use of steam, in the later half of the eighteenth century has its roots in devices that were invented earlier in the era.


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Each advancement in the industrial revolution has brought with it a certain amount of risk and hazardous activity. Before the industrial revolution, the chemical hazards we strive to manage safely today did not exist because none of these chemicals were used. As global competition increased, so did the need for using increasingly complex processes and new and exotic chemicals at different operating conditions. Demographic changes and the unplanned growth of bedroom communities around industrial belts have also created a higher risk. Notwithstanding the benefits of the industrial revolution, there has also been considerable dialogue about the risks posed by these industrial facilities. The subject of risk-benefit analysis has been discussed to some extent but is not quite formalized. Based somewhat on post-accident emotions and partly on a realization that industry can and should operate safely, there has been a groundswell of opinions from various stakeholders for structured programs to improve safety.

Government Regulations The history of safety regulations in the United States can be traced back to 1899, when the United States government issued the River & Harbor Act. This act prohibited the creation of any obstruction not authorized by Congress, to the navigable capacity of any waters of the United States except on plans authorized by the Secretary of the Army. The act was promulgated expressly to protect the nation’s waterways from excessive dumping. Subsequent to the River and Harbor Act, Congress has passed numerous laws, which impose environmental or safety regulations on businesses. In 1936, the federal government enacted the Walsh–Healy Act to establish federal safety and health standards for activities relating to federal contracts. The Walsh–Healy Act led to early research into the identification and control of occupational diseases. The ideas behind this act are the basis of many of today’s occupational health and safety regulations. During the period between 1936 and 1970, a number of other regulations were promulgated. For example, the Federal Water Pollution Control Act, the Atomic Energy Act, the Metal and Non-Metallic Mine Safety Act, and the Federal Coal Mine Health and Safety Act. Although some progress was made, these regulations were never sufficiently supported to carry out a satisfactory program. This produced relatively inconsistent and ineffective results. In 1970, Congress promulgated the Occupational Safety and Health Act (OSHAct). As a result of this landmark legislation, OSHA and the National Institute for Occupational Safety and Health (NIOSH) were established within the Department of Labor and the Department of Health and Human Services, respectively. OSHA’s mission is to ‘‘Assure so far as possible every working man and woman in the nation safe and healthful working conditions.’’ The OSHAct allows OSHA to set and enforce standards that require employers to maintain safe and healthful workplaces. NIOSH, on the other hand, does not have any regulatory or enforcement authority, but it is charged with the responsibility of training professionals and with the research and recommendation of new regulations to the Secretary of Labor.

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Process Safety/Risk Management Regulations Environmental issues affecting the public health and the environment also received widespread attention. As a result, the Environmental Protection Agency (EPA) was established in 1970 to protect the nation’s public health and environment. The EPA is responsible ‘‘to find ways to clean up and prevent pollution, ensure compliance and enforcement of environmental laws, assist states in environmental protection efforts, and scientific research and education to advance the nation’s understanding of environmental issues.’’ In 1970, the EPA promulgated the Clean Air Act, followed by amendments to the act in 1977 and 1990. The Toxic Substances Control Act (TSCA), passed in 1976, gave the EPA the ability to track and study the 75,000 industrial chemicals produced or imported to the United States. The TSCA is a federally enforced law and is not delegated to the states. Under this act, the EPA has the authority to ban the manufacture or distribution in commerce, limit the use, require labeling, or place other restrictions on chemicals that pose unreasonable risk. Asbestos, chlorofluorocarbons, and polychlorinated biphenyls are some of the chemicals regulated by the EPA under TSCA. In 1977, the International Safe Container Act established uniform structural requirements for international cargo containers designed to be transported interchangeably by sea and land carriers. In 1983, the Surface Transportation Assistance Act established protection from reprisal by employers for truckers and certain other employees in the trucking industry involved in activity related to interstate commercial motor vehicle safety and health. In December 1984, the release of 40 metric tons of methyl isocyanate from a pesticide manufacturing plant in Bhopal, India caused the deaths of over 2000 people and injuries to another 100,000 [10]. As a direct consequence of this incident, the U.S. Congress in 1986 promulgated the Emergency Planning and Community Right-to-Know Act (EPCRA). The EPCRA requires manufacturers, users, and storage facilities to keep records about quantity, use, and release of hazardous materials and make these records available for public record. The EPCRA provided pathways for better understanding of chemical hazards and called for community emergency response procedures at the local and state levels. Subsequently, the EPCRA led to the formation of Local Emergency Planning Committees (LEPCs) and State Emergency Response Commissions (SERCs). The LEPCs are voluntary organizations at the local level and are responsible for developing local emergency response plans in coordination and collaboration with local industry. The SERCs, on the other hand, are state organizations responsible for coordinating the local emergency response plans and administering state programs. EPCRA’s reporting requirements and emergency planning and notification provisions established a coordinated effort among EPA, state governors, SERCs and LEPCs, owners/operators of regulated facilities, and local fire departments. LEPCs receive chemical inventory information, analyze the hazards, and develop local emergency response plans. The LEPCs are responsible for disseminating this information to the public and serving as a focus for community awareness and action. The SERCs are appointed by the governors and consist of state emergency, environmental, and health agencies, public interest associations, and others with emergency management experience. The LEPC’s makeup is specified by law, typically consisting of the following:


• • • • •

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Representatives of elected state and local officials Law enforcement officials, civil defense workers, and firefighters First aid, health, hospital, environmental, and transportation workers Representatives of community groups and the news media Owners and operators of industrial plants and other users of chemicals, such as hospitals, farms, and small businesses

The EPCRA extended right-to-know beyond the workplace and into the community. This information has stimulated communication between industries and communities and encouraged industries to store smaller inventories of hazardous substances, discharge less, and substitute less hazardous chemicals. One major drawback of this initiative is the unfunded and voluntary nature of the LEPCs. As a result, LEPCs in many counties are marginally active or do not exist at all. In 1989, OSHA published recommended Safety and Health Program Guidelines. These voluntary guidelines identify four general elements that are critical to the development of a successful safety and health management program. These are management commitment and employee involvement, worksite analysis, hazard prevention and control, and safety and health training. In 1990, EPA analyzed chemical incidents in the early to mid-1980s and compared them to the Bhopal incident. The analysis concluded that 17 incidents released sufficient volumes of chemicals that could have been more severe than Bhopal if the weather conditions and plant location were different. Thus, the Clean Air Act Amendments of 1990 contained specific mandates requiring OSHA and EPA to establish regulations to protect workplace employees and the public and the environment, respectively. OSHA fulfilled its mandate in 1992 by promulgating the process safety management regulation. The EPA, on the other hand, promulgated the Risk Management Program regulation in 1996. The Clean Air Act Amendments of 1990 also established the Chemical Safety and Hazard Investigation Board.

State and Local Government Roles During the late 1800s and early to mid-1900s, the majority of worker-safety laws were enacted by the state and local governments and thus varied widely in their extent and enforcement. By statute, individual states have the option of seeking the delegation of most federal safety regulations. The state may request delegation from the federal government and submit a state implementation program. The state implementation program must, in content and enforcement, be, at the minimum, as stringent as the federal regulation. Twenty-three states and two U.S. territories operate their own OSHA-approved occupational safety and health programs. These ‘‘State plan States’’ are integral partners in OSHA’s mission of assuring the safety and health of the nation’s workers. They are not required to operate programs identical to those of the federal

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Process Safety/Risk Management Regulations OSHA, but they have the flexibility to operate programs that reflect their own statespecific issues and concerns, provided their programs are ‘‘at least as effective’’ as the federal OSHA program. States must set job safety and health standards that are ‘‘at least as effective as’’ comparable federal standards. Most states adopt standards identical to federal ones. States have the option to promulgate standards covering hazards not addressed by federal standards. A state must conduct inspections to enforce its standards, cover public (state and local government) employees, and operate occupational safety and health training and education programs. In addition, most states provide free on-site consultation to help employers identify and correct workplace hazards. Such consultation may be provided either under the plan or through a special agreement under Section 21(d) of the act. To gain OSHA approval for a ‘‘developmental plan,’’ the first step in the state plan process, a state must assure OSHA that within 3 years, it will have in place all the structural elements necessary for an effective occupational safety and health program. These elements include appropriate legislation, regulations and procedures for standards setting, enforcement, appeal of citations and penalties, and a sufficient number of qualified enforcement personnel. Once a state has completed and documented all its developmental steps, it is eligible for certification. Certification renders no judgment as to actual state performance, but merely attests to the structural completeness of the plan. At any time after initial plan approval, when it appears that the state is capable of independently enforcing standards, OSHA may enter into an ‘‘operational status agreement’’ with the state. This commits OSHA to suspend the exercise of discretionary federal enforcement in all or certain activities covered by the state plan. The ultimate accreditation of a state’s plan is called ‘‘final approval.’’ When OSHA grants final approval to a state under Section 18(e) of the act, it relinquishes its authority to cover occupational safety and health matters covered by the state. After at least 1 year following certification, the state becomes eligible for final approval if OSHA determines that it is providing, in actual operation, worker protection ‘‘at least as effective’’ as the protection provided by the federal program. The state also must meet 100% of the established compliance staffing levels (benchmarks) and participate in OSHA’s computerized inspection data system before OSHA can grant final approval.

Rulemaking Process Regulations by all U.S. federal agencies are developed in a similar manner, with some minor variations. As an example, the rulemaking process followed by OSHA is discussed here. The Occupational Safety and Health Administration can begin standards-setting procedures on its own initiative or in response to petitions from other parties, including the Secretary of Health and Human Services (HHS), NIOSH, state and


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local governments, any nationally recognized standards-producing organization, employer or labor representatives, or any other interested person. If OSHA determines that a specific standard is needed, any of several advisory committees may be called upon to develop specific recommendations. There are two standing committees, and ad hoc committees may be appointed to examine special areas of concern to OSHA. All advisory committees, standing or ad hoc, must have members representing management, labor, and state agencies, as well as one or more designees of the Secretary of HHS. The two standing advisory committees are as follows:

• •

National Advisory Committee on Occupational Safety and Health (NACOSH), which advises, consults with, and makes recommendations to the Secretary of HHS and to the Secretary of Labor on matters regarding administration of the act Advisory Committee on Construction Safety and Health, which advises the Secretary of Labor on formulation of construction safety and health standards and other regulations

Recommendations for standards also may come from NIOSH, established by the act as an agency of the Department of HHS. The National Institute for Occupational Safety and Health conducts research on various safety and health problems, provides technical assistance to OSHA, and recommends standards for OSHA’s adoption. While conducting its research, NIOSH may make workplace investigations, gather testimony from employers and employees, and require that employers measure and report employee exposure to potentially hazardous materials. NIOSH also may require employers to provide medical examinations and tests to determine the incidence of occupational illness among employees. When such examinations and tests are required by NIOSH for research purposes, they may be paid for by NIOSH rather than the employer. Once OSHA has developed plans to propose, amend, or revoke a standard, it publishes these intentions in the Federal Register as a ‘‘Notice of Proposed Rulemaking’’ or often as an earlier ‘‘Advance Notice of Proposed Rulemaking.’’ An ‘‘Advance Notice’’ is used, when necessary, to solicit information that can be used in drafting a proposal. The Notice of Proposed Rulemaking will include the terms of the new rule and provide a specific time (at least 30 days from the date of publication, usually 60 days or more) for the public to respond. Interested parties who submit written arguments and pertinent evidence may request a public hearing on the proposal when none has been announced in the notice. When such a hearing is requested, OSHA will schedule one, and will publish, in advance, the time and place for it in the Federal Register. After the close of the comment period and public hearing, if one is held, OSHA must publish in the Federal Register the full, final text of any standard amended or adopted and the date it becomes effective, along with an explanation of the standard and the reasons for implementing it. OSHA may also publish a determination that no standard or amendment needs to be issued. The Occupational Safety and Health Administration continually reviews its standards to keep pace with developing and changing industrial technology. Therefore, employers and employees should be aware that, just as they may petition

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Process Safety/Risk Management Regulations OSHA for the development of standards, they also may petition OSHA for modification or revocation of standards.

Technology and Research Base The National Institute for Occupational Safety and Health was established by the Occupational Safety and Health Act of 1970. NIOSH is part of the Centers for Disease Control and Prevention (CDC) and is the only federal institute responsible for conducting research and making recommendations for the prevention of workrelated illnesses and injuries. The Institute’s responsibilities include the following:

• • • • •

Investigating potentially hazardous working conditions as requested by employers or employees Evaluating hazards in the workplace, ranging from chemicals to machinery Creating and disseminating methods for preventing disease, injury, and disability Conducting research and providing scientifically valid recommendations for protecting workers Providing education and training to individuals preparing for or actively working in the field of occupational safety and health

Although NIOSH and OSHA were created by the same act of Congress, they are two distinct agencies with separate responsibilities. OSHA is in the Department of Labor and is responsible for creating and enforcing workplace safety and health regulations. NIOSH is in the Department of Health and Human Services and is a research agency. The National Institute for Occupational Safety and Health identifies the causes of work-related diseases and injuries and the potential hazards of new work technologies and practices. With this information, NIOSH determines new and effective ways to protect workers from chemicals, machinery, and hazardous working conditions. Creating new ways to prevent workplace hazards is the job of NIOSH. In 1980, the U.S. Congress created the Agency for Toxic Substances and Disease Registry (ATSDR) to implement the health-related sections of laws that protect the public from hazardous wastes and environmental spills of hazardous substances. ATSDR is charged with assessing the presence and nature of health hazards at specific sites, to help prevent or reduce further exposure and the illnesses that result from such exposures, and to expand the knowledge base about health effects from exposure to hazardous substances. In 1984, amendments to the Resource Conservation and Recovery Act of 1976 (RCRA), which provides for the management of legitimate hazardous waste storage or destruction facilities, authorized ATSDR to conduct public health assessments at these sites, when requested by the EPA, states, or individuals. ATSDR was also authorized to assist the EPA in determining which substances should be regulated and the levels at which substances may pose a threat to human health. With the passage of the Superfund Amendments and Reauthorization Act of


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1986 (SARA), ATSDR received additional responsibilities in environmental public health. This act broadened ATSDR’s responsibilities in the areas of public health assessments, establishment and maintenance of toxicological databases, information dissemination, and medical education.

The Process Safety Management Program The 14 elements of the OSHA Process Safety Management (PSM) regulation (29 CFR 1910.119) were published in the Federal Register on February 24, 1992 [11]. The objective of the regulation is to prevent or minimize the consequences of catastrophic releases of toxic, reactive, flammable, or explosive chemicals. The regulation requires a comprehensive management program: a holistic approach that integrates technologies, procedures, and management practices. The process safety management regulation applies to processes which involve certain specified chemicals at or above threshold quantities, processes which involve flammable liquids or gases on-site in one location, in quantities of 10,000 lbs. or more (subject to few exceptions), and processes which involve the manufacture of explosives and pyrotechnics. Hydrocarbon fuels, which may be excluded if used solely as a fuel, are included if the fuel is part of a process covered by this regulation. In addition, the regulation does not apply to retail facilities, oil or gas well drilling or servicing operations, or normally unoccupied remote facilities. The management system required by OSHA’s process safety management regulation envisions a holistic program with checks and balances aimed at creating multiple barriers of protection. This principle is shown in Fig. 2. The performancebased approach does not prescribe specific methods and approaches, thus giving facilities the flexibility for customizing the methods to best meet their needs and organizational structures. The process hazard analysis (PHA) is the heart of the

FIG. 2 Holistic implementation of the process safety management program.

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Process Safety/Risk Management Regulations program and impacts or interfaces with all of the other elements. However, it must also be pointed out that all elements of the program must be implemented in their entirety to get the maximum benefit and accomplish the ultimate objective (i.e., reduce the frequency and severity of chemical plant accidents). Some of the other concepts that are apparent from Fig. 2 are as follows: 1. The process safety information and incident history are important inputs to the PHA and must be compiled before the PHA. 2. Employee participation is important not only for the whole program but also provides critical information during the PHA. 3. Results of the PHA should be used in modifying and or developing operating procedures, mechanical integrity program, emergency response program, and other impacted elements. 4. Irrespective of the PHA done earlier, each process change should be evaluated by the management of change program, and, if necessary, an appropriate hazard analysis should be done. 5. Pre-start-up safety review is an essential procedure for new or modified processes. Each element of the process safety program is discussed in more detail here and the interface with other elements of the process safety management is discussed. Employee Participation This element of the regulation requires developing a written plan of action regarding employee participation, consulting with employees and their representatives on the conduct and development of other elements of process safety management required under the regulation, and providing to employees and their representatives access to process hazard analyses and to all other information required to be developed under this regulation. Process Safety Information This element of the PSM regulation requires employers to develop and maintain important information about the different processes involved. This information is intended to provide a foundation for identifying and understanding potential hazards involved in the process. The process safety information covers three different areas (i.e., chemicals, technology, and equipment). A complete listing of the process safety information that must be compiled in these three areas is shown in Table 1. This information is intended to provide a foundation for identifying and understanding potential hazards involved in the process. The information in Table 1 is essential for developing and implementing an effective process safety management program. The fundamental concept is that complete, accurate, and up-to-date process knowledge is essential for safe and profitable operations. The information contained in the first column of Table 1 should be available from the Material Safety Data Sheets (MSDS) for the hazard-


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TABLE 1 Process Safety Information Chemicals Toxicity Permissible exposure limit Physical data Reactivity data Thermal and chemical stability data Effects of mixing

Technology Block flow diagram or process flow diagram Process chemistry Maximum intended inventory Safe limits for process parameters Consequence of deviations

Equipment Design codes employed Materials of construction Piping and instrumentation Diagrams Electrical classification Ventilation system design Material and energy balances Safety systems Relief system design and design basis

ous chemicals, which are used as primary or intermediate feedstocks or are produced as products at the plant. The information contained in the second column of Table 1 pertains to the technology of the process itself. The block flow diagram can be replaced by a process flow diagram. The process chemistry information must contain the basic chemical reactions involved and a brief description of the chemistry involved. The maximum intended inventory refers to the maximum amount of any regulated chemical that may be expected to be present in the whole facility at any time. The safe limits for process parameters refer to the upper and lower bounds for the process parameters outside of which the process would be hazardous. For example, in the case of a distillation process, the upper and lower limits of the process parameters outside which the operation of the process could cause significant damage to the tower or other attached equipment would have to be stated. In this example, the process parameters for which upper and lower bounds are to be specified are temperature, pressure, composition, and flow rate. The consequence of deviation from these stated bounds must also be compiled. Safe upper and lower limits for process parameters and equipment are also necessary for calibrating instrumentation. It is important to understand the distinction between process parameter limits and equipment limits. Safe upper and lower limits for process parameters can be defined as follows:

• •

For nonreactive processes, process parameter safe limits are defined based on equipment design ratings, relief device set points, and upstream and downstream conditions of the process and/or equipment. For reactive processes, process parameter safe limits are defined based on the process chemistry information or any restricted physical conditions for the reaction as well as the criteria used for nonreactive processes.

In contrast, equipment limits are specified by the manufacturer based on materials of construction and design basis. Processes must be operated and maintained within both the process parameter and equipment limits.

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Process Safety/Risk Management Regulations The final type of information that must be compiled pertains to the equipment used in the process. The intent is to assure that all equipment used in the process meets appropriate standards and codes such as those published by the American Society of Mechanical Engineers (ASME), the American Petroleum Institute (API), the American Institute of Chemical Engineers (AIChE), the American National Standards Institute (ANSI), the American Society of Testing and Materials (ASTM), and the National Fire Protection Association (NFPA). Accepted industry practices can be used to decide which standards apply. In cases where standards do not exist, generally acceptable engineering practices can be used. The materials of construction for each equipment item and the design codes employed can be compiled as separate lists or may be listed in the Piping and Instrumentation Diagram (P&ID). The P& IDs must represent the facility exactly as it exists with flanges, valves, and all other connections shown. The different electrical classifications must also be compiled for different parts of the facility. A simple plot plan showing the different areas of electrical classification would be considered to be in compliance with the regulation. Information must also be compiled on any ventilation system. This information would indicate the areas in the facility that are ventilated and the nature of ventilation. A listing of all safety systems must also be compiled that are available to the workers. This listing should include any and all equipment that is available for protection of the workers from any hazard or emergency. Information should also be available on the location of these safety systems and the procedures to use these systems. As is apparent from the foregoing discussion, compilation of the process safety information database represents a major challenge. This is complicated even more because the process safety information must also be kept up-to-date and accurate and made accessible to employees. Many plants have therefore implemented or are in the process of implementing electronic data management systems to manage, access, and use these data. The completion and accuracy of process safety information is crucial to the implementation of other PSM elements, including PHAs and mechanical integrity. The PSM regulation requires that process equipment should comply with generally accepted engineering practices. It is therefore not only important to compile all equipment information but also to ensure that it complies with consensus standards. OSHA is recognizing that there are consensus standards for design and fabrication, installation, maintenance procedures, and inspection and testing. Therefore, equipment constructed in accordance with codes, standards, or practices no longer in use should be evaluated to ensure that they are compatible with existing standards. For example, a multilayered vessel built years ago should be evaluated to ensure that it complies with today’s engineering standards.

Process Hazards Analysis This element of the PSM regulation requires facilities to perform a PHA. The PHA must address the hazards of the process, previous hazardous incidents, engineering and administrative controls, the consequences of the failure of engineering and administrative controls, human factors, and an evaluation of effects of failure of controls on employees. This element requires that the PHA be performed by one or more of the following methods or any other equivalent method:


• • • • • •

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What If Checklist What If/checklist Hazard and operability (HAZOP) studies Failure modes and effects analysis (FMEA) Fault-tree analysis

The regulation suggests a performance-oriented requirement with respect to the PHA so that the facility has the flexibility to choose the type of analysis that will best address a particular process. PHAs may not be performed unless complete Process Safety Information is available for the process. Process hazard analyses and the results of PHAs can and will impact the development, implementation, and practice of other elements of the PSM regulation. A PHA would help facilities identify hazards and ways to address them. For example, a 1989 explosion and fire at a facility in Baton Rouge, Louisiana led to a partial loss of pressure, power, and fire water because the power, steam, and water lines were colocated with the lines carrying flammable gases [12]. The losses complicated and prolonged the process of responding to the release, thereby increasing the damage caused by the release. Similar problems occurred at a facility in Norco, Louisiana, where an explosion led to the loss of all utilities. A thorough and properly done PHA should identify these types of potential hazards and allow facilities to determine how to mitigate the problems. PHAs also identify situations where major accidents due to control failure (e.g., pressure gauges, overfill alarms) could be prevented by redundant or backup control or by frequent maintenance and inspection practices. Many other elements of the PSM program should flow from, or at least be revised based on, the results of the PHA. Existing standard operating procedures, training and maintenance programs, and pre-start-up safety reviews may need to be revised to reflect changes in either practices or equipment that derive from the PHA. The PHA may help define critical equipment that require preventive maintenance, inspection, and testing programs. It may also help a facility focus its emergency response programs on the most likely and most serious release scenarios. The PSM regulation also requires that the PHAs be updated and revalidated, based on their completion date. First, the PSM regulation requires that the process hazard analysis shall be updated and revalidated. Thus, the whole PHA should not only be examined to verify that the PHA is consistent with the current process, but it also should be analyzed to examine if the original analysis is valid. Second, the PSM regulation requires that the updating and revalidating be done by a team similar in qualifications to the team that conducted the original PHA. The intent of a PHA revalidation is to ensure that the PHA team evaluates the previous PHA, examines the extent of any changes that might have occurred since the PHA was implemented (or last reviewed), and decides what work is needed to make the PHA current. Philosophically, in an ideal process safety management system, the following process safety concepts should work as checks and balances to accomplish the PSM program objectives (i.e., minimize the frequency and consequences of catastrophic releases):

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• • •

Initial process hazards analysis: Identify process hazards and develop mitigation techniques (technology, equipment, and procedures). Management of change: Assess the safety and health impacts of process changes and ensure that process changes are analyzed (e.g., HAZOPed). Revalidation of process hazards analysis: Update and revalidate the PHAs to assure that the process hazard analysis is consistent with the current process.

Operating Procedures The operating procedures must be in writing and provide clear instructions for safely operating processes, must include steps for each operating phase, operating limits, safety and health considerations, and safety systems. Procedures must be readily accessible to employees, must be reviewed as often as necessary to assure they are up-to-date, and must cover special circumstances such as lockout/tagout and confined space entry. The employer must certify annually that the operating procedures are current and accurate. The synergism and commonality of operating procedures to maintenance procedures is in safe work practices. These safe work practices include lockout/tagout, confined space entry, opening of process equipment or piping (i.e., hot tapping), and control of entrance into the battery limits. Even though the operations department is involved with all of these practices, maintenance also plays a very vital role in ensuring that these procedures are followed during all maintenance tasks. Many incidents have resulted from inadequate safe work practices or a failure to follow procedures when they exist.

Training The regulation requires that facilities certify that employees responsible for operating the facility have successfully completed (including means to verify understanding) the required training. The training must cover specific safety and health hazards, emergency operations, and safe work practices. Initial training must occur before assignment. Refresher training must be provided at least every 3 years. Even though this element of the PSM regulation pertains to operations staff only, it is important to remember that operations and maintenance training should be coordinated.

Contractors The PSM regulation identifies responsibilities of the employer regarding contractors involved in maintenance, repair, turnaround, major renovation or specialty work, on or near covered processes. The host employer is required to consider safety records in selecting contractors, inform contractors of potential process hazards, explain the facility’s emergency action plan, develop safe work practices for contractors in process areas, periodically evaluate contractor safety performance, and maintain an injury/illness log for contractors working in process areas. In addition, the contract employer is required to train their employees in safe work


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practices and document that training, assure that employees know about potential process hazards and the host employer’s emergency action plan, assure that employees follow safety rules of facility, and advise host employer of hazards contract work itself poses or hazards identified by contract employees. In the contractor paragraph, OSHA has used a belt and suspender approach. Both the host employer and contract employer have specific responsibilities that they must fulfill. The need for flexibility, quick turnarounds, and specialized services is the main reason why process plants are contracting out increasingly significant portion of their daily work, particularly maintenance work to contractors. Contractors, in general, receive less training and often perform more hazardous tasks in process plants as compared to direct-hire workers [13].

Pre-Start-up Safety Review This element of the PSM regulation requires a pre-start-up safety review of all new and modified facilities to confirm integrity of equipment, to assure that appropriate safety, operating, maintenance, and emergency procedures are in place, and to verify that a process hazard analysis has been performed. Modified facilities for this purpose are defined as those for which the modification required a change in the process safety information. Usually, changes occur during maintenance and, therefore, maintenance personnel should be well versed in pre-start-up safety review procedures. Maintenance should ensure that all necessary procedures have been completed prior to start-up.

Mechanical Integrity This element of the PSM regulation mandates written procedures, training for process maintenance employees, and inspection and testing for process equipment, including pressure vessels and storage tanks, piping systems, relief and vent systems and devices, emergency shutdown systems, pumps, and controls such as monitoring devices, sensors, alarms, and interlocks. PSM calls for correction of equipment deficiencies and assurance that new equipment and maintenance materials and spare parts are suitable for the process and properly installed.

Hot Work Permit This element of the PSM regulation mandates a permit system for hot work operations conducted on or near a covered process. The purpose of this element of the regulation is to assure that the employer is aware of the hot work being performed and that appropriate safety precautions have been taken prior to beginning the work. Because welding shops authorized by the employer are locations specifically designated and suited for hot work operations, the regulation does not require a permit for hot work in these locations. Additionally, hot work permits are not required in cases where the employer or an individual to whom the employer has assigned the authority to grant hot work permits is present while the hot work is performed.

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Process Safety/Risk Management Regulations Management of Change This element of the regulation specifies a written program to manage changes in chemicals, technology, equipment, and procedures which addresses the technical basis for the change, impact of the change on safety and health, modification to operating procedures, time period necessary for the change, and authorization requirements for the change. The regulation requires employers to notify and train affected employees and update process safety information and operating procedures as necessary. Incident Investigation This element of the regulation requires employers to investigate as soon as possible (but no later than 48 hrs) incidents which did result or could have resulted in catastrophic releases of covered chemicals. The regulation calls for an investigation team, including at least one person knowledgeable in the process (a contractor employee, if appropriate), to develop a written report of the incident. Employers must address and document their response to report findings and recommendations and review findings with affected employees and contractor employees. Reports must be retained for 5 years. Emergency Planning and Response This element requires employers to develop and implement an emergency action plan according to the requirements of 29 CFR 1910.38(a) and 29 CFR 1910.120(a), (p), and (q). Compliance Audits This element of the regulation requires employers to certify that they have evaluated compliance with process safety requirements every 3 years and specifies retention of the audit report findings and the employer’s response. The employer must retain the two most recent audits. Trade Secrets Similar to the trade secret provisions of the hazard communication regulation, the PSM regulation also requires information to be available to employees from the process hazard analyses and other documents required by the regulation. The regulation permits employers to enter into confidentiality agreements to prevent disclosure of trade secrets. As is apparent from the foregoing discussion, the process safety management regulation requires a systems approach for managing safety. Segments of the hazardous chemicals industry have for sometime practiced some or all of the required programs. The promulgation of the regulation formalized the requirements and


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established minimum criteria. This is both good and bad. The regulation now requires everyone to establish the management systems and apply the technologies needed to comply with the regulation. However, because of the same reason, there is a tendency to look for ‘‘paper compliance’’ as compared to making real improvements in safety programs and technologies.

The Risk Management Program In 1996, EPA promulgated the regulation for Risk Management Programs for Chemical Accident Release Prevention (40 CFR 68). This federal regulation was mandated by Section 112(r) of the Clean Air Act Amendments of 1990. The regulation requires regulated facilities to develop and implement appropriate risk management programs to minimize the frequency and severity of chemical plant accidents. In keeping with regulatory trends, the EPA required a performance-based approach toward compliance with the risk management program regulation. The eligibility criteria and requirements for the three different program levels are given in Table 2. The EPA regulation also requires regulated facilities to develop a Risk Management Plan (RMP). The RMP includes a description of the hazard assessment, prevention program, and the emergency response program. Facilities submit the RMP to the EPA and, subsequently, it is made available to governmental agencies, the state emergency response commission, and the local emergency planning committees and is communicated to the public. The RMP regulation defines the worst-case release as the release of the largest quantity of a regulated substance from a vessel or process-line failure, including administrative controls and passive mitigation that limit the total quantity involved or release rate. For gases, the worst-case release scenario assumes the quantity is released in 10 min. For liquids, the scenario assumes an instantaneous spill and that the release rate to the air is the volatilization rate from a pool 1 cm deep unless passive mitigation systems contain the substance in a smaller area. For flammables, the scenario assumes an instantaneous release and a vapor cloud explosion using a 10% yield factor. For alternative scenarios (note: the EPA used the term alternative scenario as compared to the term more-likely scenario used earlier in the proposed regulation), facilities may take credit for both passive and active mitigation systems. Appendix A of the final regulation lists endpoints for toxic substances to be used in worst-case and alternative scenario assessment. The toxic endpoints are based on ERPG-2 (Emergency Response Planning Guidelines—Level 2) or level of concern data compiled by EPA. The flammable endpoints represent vapor cloud explosion distances based on overpressure of 1 psi or radiant heat distances based on exposure to 5 kW/m2 for 40 s.

Impact on Process Industry In general, the impact of any regulation on any part of the regulated community can be related to several factors which include, but are not limited to the following:

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Process Safety/Risk Management Regulations TABLE 2 Eligibility Criteria and Compliance Requirements for Different Program Levels; EPA’s Risk Management Program Regulation Program 1 No off-site accident history

Program 2

Program Eligibility Criteria Process not eligible for pro- Process is subject to OSHA gram 1 or 3 PSM (29 CFR 1910.119)

No public receptors in worst-case circle Emergency response coordinated with local responder

Hazard Assessment Worst-case analysis 5-Year accident history Certify no additional steps needed

Process is SIC code 2611, 2812, 2819, 2821, 2865, 2869, 2873, 2879, or 2911 Program Requirement Hazard Assessment Hazard Assessment Worst-case analysis Worst-case analysis Alternative releases Alternative releases 5-Year accident history 5-Year accident history Management Program Document management system Prevention Program Safety information Hazard review Operating procedures Training Maintenance Incident investigation Compliance audit

Emergency Response Program Develop plan and program

• • •

Program 3

Management Program Document management system Prevention Program Process safety information Process hazard analysis Operating procedures Training Mechanical integrity Incident investigation Compliance audit Management of change Pre-start-up safety review Contractors Employee participation Hot work permits Emergency Response Program Develop plan and program

The degree to which a particular entity or entities is already attempting to reach a particular level of achievement The additional cost associated with regulatory overhead not directly associated with the goals of the program (e.g., the cost of record keeping) The degree to which already successful programs are compromised by the regulation


• • • • • •

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The technical feasibility of compliance The impact on compliance with other regulations The emphasis on compliance rather than impact The ability to maintain a profitable operation The ability to absorb the associated overhead cost when offshore competition does not have those costs The degree to which enforcement and other emphasis by the regulating authority promotes real improvement as opposed to ‘‘paperwork’’ compliance

It must also be noted that painting all industry or even one segment of industry with one brush is very inappropriate. Many companies and/or groups of companies are and have been for many years investing considerable time and resources into process safety and risk management prior to any regulatory requirement to do so. Also, many companies and industry groups have initiated voluntary programs which go further than regulations require. In some cases, these efforts have been in partnership with government and nongovernment groups and in other cases they have been independent of government participation.

Compliance Programs At first glance, it would seem that ‘‘compliance’’ must imply compliance with regulations. However, there are many examples of responsible industry doing something because it was ‘‘the right thing to do.’’ Following the Texas City disaster in 1949, the industries on the Houston Ship Channel voluntarily formed Channel Industries Mutual Aid (CIMA) in order to help each other and the public during emergencies. There were no regulatory drivers. Responsible chemical companies have been responding to transportation emergencies involving their products for at least 50 years. Safety programs for worker protection were in place years before OSHA. For example, dating back to the 1960s and 1970s, Dow Chemical Company has had an excellent reactive chemicals program. A positive result of regulatory action has been that companies who have not voluntarily invested in these kinds of programs are forced to make the investment. Unfortunately, when compliance with regulation becomes the issue, administrative overhead may increase significantly. Furthermore, a one-size-fits-all approach to regulation and compliance may force a company to abandon a program that has proven successful over time simply to meet a regulation. It is very difficult to explain to management and workers that a highly successful, popular program has to be changed or eliminated because of a rigidly written regulation. Many companies had employed some form of lockout/tagout procedure for years that worked. The one size fits all regulatory mentality caused those programs to be changed at great cost and effort. It can be argued that the cost was not justified by the result. The changed requirements also caused confusion in the case of some workers, which might actually have increased the risk. The good news is that those same responsible companies extend their policies to their facilities wherever they may be. For example, members of the American Chemistry Council apply the Responsible Care Code internationally.

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Process Safety/Risk Management Regulations Finally, when the ultimate goal becomes compliance rather than results, it is possible to be in technical compliance and not in functional compliance. For example, several persons both in and out of government during work on the National Response Team’s integrated contingency plan effort commented that in many cases facility and vessel owners hired contractors to produce response plans under the Oil Pollution Act of 1990 that were judged by government contractors to meet all planning requirements but were not useful for response. In other words, compliance, not planning and response, was the goal.

Enforcement Enforcement can be a two-edged sword. As mentioned earlier, compliance issues can sometimes outweigh the issues of functionality or statutory and regulatory intent. Very often, with both government and industry, the issue becomes ‘‘did you meet the requirements’’ not ‘‘have you increased or reduced the risk.’’ At the same time, it must also be said that in some cases, the threat of enforcement is the only leverage regulators have. Sometimes, however, the threat of enforcement may actually serve to increase overall risk. There is currently a great deal of concern that efforts to mix enforcement and response at the incident command level in oil spills may discourage the cooperative spirit necessary for an effective response. Overall, the availability of enforcement has probably had a positive impact on process safety and risk management. The threat of enforcement has probably had a chilling effect on cooperative efforts that could possibly make an even larger impact.

Operational Issues Operationally, regulations have had a net positive impact in that they have forced many companies not already committed to environmental and safety excellence to operate better. In many cases, this improvement has actually been transferred to the bottom line. Improved operating discipline actually will increase quality, productivity, morale, and product yield. There are occasional situations, however, that require procedures that may cause significant problems. For example, there has never been a catastrophic explosion of an underground fuel tank. That is why many fire codes prohibit aboveground tanks at retail facilities. The underground storage tank regulations have made it very expensive to maintain and install underground tanks. In this case, there is an obvious conflict between two competing interests. Some air regulations require inspection of the seals on floating roof tanks. It is very difficult to clear those tanks well enough to eliminate emissions during the inspection process. In this case, one regulation may result in violation of another or an actual increase in emissions. Elimination of chlorofluorocarbons has caused a concurrent elimination of Halon-type fire-extinguishing systems. Another example is the requirement to test marine facility foam systems by actually flowing foam. These tests have a high likelihood of discharging some foam to the water—a violation of the Clean Water Act.


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Management Systems for Compliance Management systems for regulatory compliance have evolved over a period of time. In the early stages, there was a tendency to manage process safety as a separate function. With experience, industry has learned that is not the most efficient or effective approach. In order to gain the most advantage, safety must be integrated into the plant’s everyday activities through each phase of the plant’s life cycle. The following provides a discussion of some example components of the management systems necessary for the effective implementation of process safety and risk management programs.

Integrated Systems Versus Compartmentalized Systems Management systems for compliance must be integrated and cannot be compartmentalized. First, although statutes and the accompanying regulations are often conceived and written as separate entities, they often have impacts on other areas that may or may not be intentional. One example can be found in the list of chemicals regulated under the Clean Air Act Amendments that included ethylene glycol. Under the reportable quantity (RQ) regulations of the Superfund Amendments and Reauthorization Act, ethylene glycol defaulted to a 1-lb RQ. This had the effect of creating a large number of unnecessary and wasteful emergency release reports from people who understood the regulations and in one sense made criminals out of almost everyone who had a radiator leak. Only an integrated system could respond to that quirk. The real thrust of integrated management systems must be to include all safety and environmental concerns in the entire process from conception to design to construction and commissioning to start-up and operation to shutdown and demolition. Only this approach can insure compliance, but more importantly safety and environmental excellence. Individuals in the organization must understand their respective roles. Therefore, the management system for compliance should be integrated vertically within the organization and horizontally across regulatory regimes.

Process Design It is critical that companies have a system in place that assures compliance from the time of conception to eventual start-up of any process. There are multiple regulations that impact every facet of a new process. The process design must look at air emissions and permitting, wastewater and stormwater handling, waste handling, generation and reduction, Toxic Substance Control Act implications, Risk Management Plan implications, and other potential areas of concern. Worker safety, including repetitive motion injuries, should be an issue from the start. Soft issues such as public perception must be considered. All of these issues also apply to modifications to existing facilities. They may also apply to changes in operating procedures and other ‘‘soft’’ changes. For all of these reasons and for OSHA compliance, a well-conceived management of change (MOC) system must be integrated into the culture.

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Process Safety/Risk Management Regulations Very subtle, often seemingly innocuous changes can have significant regulatory implications. An operator of a water-treatment facility may have several 1-ton chlorine cylinders on site but never have more than one connected to the process. That operator may decide that the greatest risk occurs when changing the cylinder out. The obvious solution would be to manifold several cylinders together. That change would then make the operator subject to the provisions of Section 112r of the Clean Air Act.

Equipment Maintenance For all of the reasons listed earlier, equipment maintenance must be part of the management system. It is possible that regulatory issues may actually discourage a good maintenance program. Many maintenance activities, especially in older plants, require opening processes to the atmosphere in order to make the equipment safe to work. Air regulations may prohibit, or at least discourage, this activity. There is no way of quantifying how often needed maintenance is delayed or canceled because of environmental concerns. Because of this, a relatively small emission or environmental upset resulting from maintenance was avoided at the cost of an eventual catastrophic event.

Prevention and Reduction Responsible industry has always had prevention and reduction programs. As stated earlier, regulations have largely served to ‘‘level the playing field’’ between responsible and irresponsible industry. As early as the 1960s, one chemical company executive stated on the ‘‘Today’’ show that his company was in the business of making salable product out of raw materials, and if a molecule left the plant as waste to the environment, it represented lost profit. Further, in recent years, all companies have begun to recognize that they must be good neighbors by perception as well as by reality. Initiatives such as Responsible Care illustrate this recognition.

Response Mechanism Again, responsible companies have had response systems in place for many years. Several things have occurred to alter and, in most cases, improve those systems. The OSHA PSM regulation and the Hazardous Waste Operations regulation (OSHA 1910.120) have elevated the training requirements for responders. Although experienced responders already had most of the required training and/or experience, the regulations created minimums that have been very effective. Planning requirements under a multitude of federal and state regulations have, however resulted in considerable duplication. PSM, RMP, Community Right to Know, OPA90, RCRA, and others all mandate response planning. The vessel response plan requirements proposed by the U.S. Coast Guard will add still further requirements. This duplicative regulatory system resulted in numerous plans designed solely to meet compliance requirements and often had little to do with response. The good news is that the effort of the National Response Team to publish the


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Integrated Contingency Planning Guidance has relieved much of that problem. In the case of OPA90 in particular, oil spill response is much better than it was during the Exxon Valdez spill. The most lasting improvement has been in the response partnerships formed between industry and government at all levels.

Risk Communication For many years, much of industry did a fairly poor job of telling people about potential risks. This has been significantly improved by Community Right to Know, 112r of the Clean Air Act, and voluntary programs such as Responsible Care. Much remains to be done in this area. Industry and government tend to respond to public perception of risk without doing a good job of education about true risk. Very often, uninformed public outcry has forced questionable, if not totally incorrect, reaction on the part of industry and government. All too often, groups with hidden agendas have aggravated this process. Fears about dioxins that have not been scientifically confirmed have resulted in actions like Times Beach; this is in spite of the record of Seveso in which there are still no documented longterm effects of the exposure. The challenge is creating an interactive dialogue with the public to a point where risk perception is based not only on emotion but also on some level of scientific reality. Zero risk is a myth. However, zero negative impact is a vision for which industry and government must strive.

Small Business Issues In the United States, a significant portion of the economy consists of small and midsized companies. We must remember that an accident from such a small facility has the likelihood of severe consequences and can damage the whole industry ‘‘license to operate’’ just like an accident in a plant operated by a large multinational company. However, the large facility probably has resources, training, and equipment either to prevent the accident in the first place or respond to the consequences if it does occur. On the other hand, the small facility probably lacks awareness, training, information, and other things. A recent accident in the United States resulted in multiple fatalities and total destruction of a small plant. Preliminary investigations highlight the lack of reactive chemicals knowledge among the plant personnel. The challenge therefore is to develop collaborative efforts with government agencies (both state and federal), professional and trade organizations, and industry for safety programs aimed at improving safety in the small and mid-sized companies.

Future Developments In an increasingly global economy in which the developed nations will compete even more strongly with the less developed nations, several things must happen.

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Process Safety/Risk Management Regulations First, multinational companies must insist that the best practices of worker, public, and environmental protection be followed throughout the world. Product stewardship is one effective method for ensuring the implementation of the same standards at small and medium-sized businesses in the United States as well as in facilities overseas. Governments should also promote those best practices. Industries and governments not doing so should face sanctions in the international marketplace. It is imperative that those companies trying to do what is right are not penalized by unfair competition. Partnerships between key players in this arena should be encouraged and rewarded. Examples include Integrated Contingency Planning and the OSHA Voluntary Protection Program. The public should be represented in these programs. All of the stakeholders involved in these issues must work to identify and eliminate barriers to improvement. These barriers and issues include, but are not limited to the following:

• • • • •

Tort regimes that discourage sharing of lessons learned and near-miss information Regulations and requirements that are duplicative or not based on science Public policy that results in ‘‘knee-jerk’’ statutory and regulatory response to single events Companies and industry groups that are willing to conduct meaningful dialogue with stakeholders Most importantly, a willingness to accept and work toward a vision in which no facility has a negative impact on its workers, the public, or the environment.

Future developments in the United States with regard to process safety and risk management programs may quite likely be based on risk–benefit analyses. There is also number of efforts underway to develop stakeholder dialogue and arrive at consensus opinions regarding national safety goals and targeted improvements in safety performance. It is quite clear that the need to operate safely is recognized as a competitive advantage and a positive contributor to the bottom line. The regulatory regime and requirements will also keep changing as more information becomes available. Thus, industrial programs and practices will have to keep pace with the changing clime and consensus standards and targeted safety goals.

Summary and Conclusions Safety regulations in the United States have mirrored the industrial revolution. The industrial revolution brought prosperity along with the use of hazardous processes. As our understanding of the hazards associated with these processes developed, procedures and practices were put in place to limit or eliminate the damage. Government programs and industry initiatives spurred improvements in the science and technology needed for the recognition of hazards and associated risks. Management systems have been developed to implement safety programs and industry practices.


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Government regulations continue to be a significant driver for safety programs. As such, one of the main objectives of the management systems is to ensure compliance. However, it is also quite clear that profitability is directly related to safety and loss prevention. Thus, the management systems for safety are intricately tied into the operational management. It is also quite apparent that government regulations alone cannot accomplish our ultimate safety goals. Safety must be tied into profitability and business objectives. Also, the use of risk assessments and risk– benefit analysis will play a significant role in the future safety programs.

References

1. U.S. Environmental Protection Agency, Accidental Release Information Program, EPA, Washington, DC, 1989. 2. M. Mannan, A. Akgerman, R. G. Anthony, R. Darby, P. T. Eubank, and K. R. Hall, ‘‘New Challenges in Chemical Engineering: Integrating Process Safety into Chemical Engineering Education and Research,’’ Chem. Eng. Ed., 33(3), 198–209 (1999). 3. U.S. Environmental Protection Agency, A Review of Federal Authorities for Hazardous Materials Accident Safety: Report to Congress Section 112(r)(10) Clean Air Act As Amended, EPA, Washington, DC, 1993. 4. National Environmental Law Center and United States Public Interest Research Group, Accidents Do Happen: Toxic Chemical Accident Patterns in the United States, United States Public Interest Research Group, Washington, DC, 1994. 5. Code of Federal Register, 40 CFR 372, U.S. Environmental Protection Agency, Toxic Chemical Release Reporting, Community Right-to-Know, FDA, Washington, DC, June 26, 1991. 6. T. A. Kletz, Lessons from Disaster, Gulf Publishing Company, Houston, TX, 1993. 7. Skandia International, Bleve—The Tragedy of San Juanico, Skandia International, Stockholm, 1985. 8. W. D. Cullen, The Public Inquiry into the Piper Alpha Disaster, Her Majesty’s Stationery Office, London, 1990. 9. U.S. Department of Labor, The Phillips 66 Company Houston Chemical Complex Explosion and Fire, U.S. Department of Labor, Washington, DC, 1990. 10. P. Shrivastava, Bhopal—Anatomy of a Crisis, Ballinger, Cambridge, MA, 1987. 11. ‘‘Process Safety Management of Highly Hazardous Chemicals; Explosives and Blasting Agents; Final Rule, 29 CFR Part 1910, Department of Labor, Occupational Safety and Health Administration, Washington, DC, February 24, 1992,’’ Fed. Reg., 57(36), 6356–6417 (1992). 12. Risk Management Programs for Chemical Accidental Release Prevention; Proposed Rule; 40 CFR Part 68, Environmental Protection Agency, Washington, DC, October 20, 1993, Fed. Reg., 58(201), 54,190–54,219 (1993). 13. John Gray Institute, Managing Workplace Safety and Health: The Case of Contract Labor in the U.S. Petrochemical Industry, John Gray Institute, Lamar University System, 1991. M. SAM MANNAN JIM MAKRIS H. JAMES OVERMAN

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Recent Advances in Biomaterials


A vast number of materials (collectively termed biomaterials) are used in clinical applications, from drug delivery (capsules, transdermal patches) to hip replacements to materials in hemodialysis machines to heart valves. These materials have had a tremendous impact on patients’ lives, allowing many injured or ill people to lead a near-normal life in many instances. As one would imagine, these biomaterials represent a significant health care cost. The annual sales of medical devices, diagnostics, and pharmaceutical products using biomaterials in some capacity exceed $100 billion per year in the United States alone. With new materials and devices being developed to treat disease and injury, the market for biomaterials is likely to increase significantly in the next decade. Specifically, a biomaterial can be classified as a material such as a polymer, metal, or ceramic that is in intimate contact with a biological environment such as blood, tissue, or individual cells. The biocompatibility of a biomaterial is defined as the ability of a material to function with a specific host response. For example, a blood contacting material should not cause thrombosis or complement activation, and a hip implant should be load-bearing and non destructive to the surrounding tissue. The study of biomaterials is a multidisciplinary endeavor, integrating knowledge from fields such as engineering, chemistry, biology, and medicine. Because of the high impact of these materials on patient care, this area has been an intensely researched for the last 20 years. Although current biomaterials have had a large impact in medicine, they consist mainly of ‘‘off-the-shelf ’’ materials not originally designed for clinical use. As a result, complications have and continue to occur. For example, leakage, hardening, and rupture of silicone gel breast prostheses have lead to billions of dollars in litigation related to either real or perceived patient health problems attributed to the material. The early use of cellulose in hemodialysis tubing resulted in anaphylactic shock and death in several patients. As a result of these incidents and numerous others, many materials manufacturers no longer permit the clinical use of their materials; thus, device manufacturers have feared a shortage of materials. To help alleviate any future shortage of biomaterials, President Clinton signed the Biomaterials Access Assurance Act of 1998 to limit product liability in the biomaterials area. To eliminate many of the complications and expand the use of biomaterials in medicine, research has focused on developing materials specifically designed for the clinic. These materials are being designed using an advanced knowledge of cell biology and tissue–material interactions. Many of these new materials represent chemical and physical modifications of existing materials or are entirely novel in their chemical structure. Here, we will focus on new material development in three high-impact areas of biomaterials research: cell–biomaterial interactions, tissue engineering, and drug delivery. This discussion is by no means exhaustive, as many significant advances have also been made in other research areas that use biomaterials.


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Cell–Biomaterial Interactions Because the lack of biocompatibility of most conventional biomaterials can be attributed to adverse cell–material interactions, a large amount of research has been focused on understanding and controlling what happens at the cell–material and tissue–material interfaces. Most mammalian cells are anchorage dependent; that is, they must adhere to a substrate to grow and differentiate. Adhesion to a surface is mediated by protein adsorption onto the surface, followed by cell– receptor interactions with the adsorbed proteins. As the adherent cells grow and proliferate, they produce their own extracellular matrix (ECM) on which to adhere. Surfaces that minimize protein adsorption, such as those modified with poly (ethylene glycol), also minimize cell adhesion and spreading [1]. Cells on these surfaces are sparse in number, very loosely attached, and appear rounded. In addition to purely thermodynamic considerations such as protein adsorption, cell adhesion is strongly influenced by receptor–ligand interactions between integrins on the cell surface and ECM proteins adsorbed on the substrate. These proteins, examples of which include laminin, fibronectin, and collagen, contain peptide sequences, such as RGD (R ⫽ arginine, G ⫽ glycine, D ⫽ aspartic acid), specific for integrin binding. Materials modified with RGD peptide sequences alone are capable of integrin binding and thus permit cell attachment and growth [2]. Zhang et al. have used the RAD (A ⫽ alanine) sequence to promote cell adhesion to oligopeptides terminated with cysteine residues [3]. Via sulfhydryl interactions with a gold-coated substrate, these oligopeptides can form self-assembled monolayers (SAMs) on the surface of the substrate and thus create a good adhesion substrate for a variety of cell types. In addition to integrin–cell adhesion, cells such as hepatocytes possess receptors for asialoglycoprotein receptors for β-d-galactose residues present in the ECM [4,5]. Griffith and colleagues took advantage of these unique interactions to create an adhesion surface for hepatocytes. Starting with poly (ethylene glycol) or PEG star polymers (PEG chains radiating from a divinylbenzene core), they added d-galactose moieties to the hydroxy terminus of PEG and cross-linked the gels using electron beam irradiation. Hepatocytes readily adhered to these surfaces even though the ligand density was an order of magnitude lower than in similar galactose-modified polyacrylamide gels. The reason for this strong adhesion was that the flexibility of the PEG chains allowed necessary multiple ligand–receptor interactions to occur, even though the ligand density was low. In the past few years, Stupp and colleagues [6–8] have developed a novel class of block copolymers that self-assemble into highly ordered supramolecular structures on surfaces. These materials were diblock polymers termed ‘‘rod–coil’’ and are composed of a rigid rod segment of biphenyl esters and a flexible segment (the coil) composed of a block copolymer of polystyrene and polyisoprene (Figs. 1 and 2). When cast on surfaces, these polymers self-assembled to form identical nanometerscale clusters. The self-organization of the polymers on a surface was based on the crystallization of the biphenyl ester rod segments, as can be seen in the base of the structure depicted in Fig. 2. Molecular modeling studies suggested that these aggregates were essentially supramolecular ‘‘mushrooms’’ with the ‘‘stem’’ consisting of

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FIG. 1 Structure of self-assembling rod–coil polymer. (Adapted from Ref. 6.)

the crystallized rod segments and the ‘‘head’’ comprised of the coil segments. When annealed at high temperature, the isoprene segments cross-linked to form films stable in a variety of solvents. Because of the highly ordered and stable nature of these clusters, they hold great potential in the design of new biomaterials. Microfabrication techniques such as photolithography and microcontact printing have also been used to create surfaces with unique topographies and spatially defined chemistries. Such surfaces have had a profound influence on cell–material interactions. Researchers at Cornell, for example, have developed micromachined probes that show minimal adverse tissue response when implanted in the cerebral cortices of rats [2]. Others have shown that photolithography can define regions of cell adhesion, controlling both cell position and migration [9]. Microcontact printing or µ-Cp, a soft lithography technique, has received extensive attention as a method of both chemical patterning [10] and protein patterning surfaces [11] to control cell adhesion and migration. As described earlier, Whitesides and Lauffenberger used oligopeptide SAMs to create surfaces for cell adhesion [3]. These peptides can be patterned by µ-Cp, resulting in the spatial control of cell adhesion and migration and creating unique geometries suitable for

FIG. 2 Molecular graphics of a self-assembled rod–coil polymer structure containing nine layers or monomer units of isoprene. (From Ref. 6.)


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the study of cell–cell interactions on these surfaces. Earlier studies demonstrated that there are geometric limits to the extent cells can be confined without causing apoptosis (programmed cell death) [12–14]. Besides µ-Cp, Takayama et al. have also used the low-mixing laminar-flow conditions in micromachined channels to pattern cocultures of eukaryotic cells on a single surface [15]. Going beyond SAMs, a four-step soft lithographic process based on microcontact printing (µ-Cp) of organic monolayers, hyperbranched polymer grafting, and subsequent polymer functionalization has resulted in polymer patterns that direct the grown of mammalian cells such as IC-21 murine peritoneal macrophages [16], human umbilical vein endothelial cells, and murine hepatocytes. The functional units on these surfaces were three-dimensional cell ‘‘corrals’’ that have walls 50 nm in height and lateral dimensions on the order of 60 µm. The corrals have hydrophobic, methyl-terminated n-alkanethiol bottoms, which promote cell adhesion, and walls consisting of hydrophilic poly(acrylic acid)/poly(ethylene glycol) (PAA/PEG) layered nanocomposites that inhibit cell growth. Cells seeded on patterned surfaces adhere and grow within the corrals, but they do not span the PAA/ PEG corral walls (Fig. 3). Cell viability studies indicate that cells remain viable

FIG. 3 Murine macrophages cultured on a micropatterned surface of poly(acrylic acid)/poly(ethylene glycol). (From Ref. 16.)

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Recent Advances in Biomaterials on the patterned surfaces for up to 21 days, and microscopy studies demonstrate that cell growth and spreading does not occur outside of the corral boundaries. Other significant insights have been made in understanding tissue–material interactions, particularly in influencing the foreign-body response. When a foreign body such as a medical implant is introduced into a host, the natural tendency of the surrounding tissue is to degrade or extrude the implant. If the host cannot eliminate the foreign body, a chronic inflammatory reaction results and the object is encapsulated in fibrous tissue. This capsule poses a difficult problem in the development of cell-based therapeutic devices and engineered tissues [17]. It presents a mass-transfer barrier and therefore limits the concentration of nutrients and oxygen reaching the transplanted cells. As a result, these cells frequently die from hypoxia or lack of nutrients, particularly cells that are highly metabolically active (e.g., pancreatic islets and hepatocytes) [18]. The ability of the capsule to limit mass transfer has been demonstrated in several studies. In one early study, fibrous capsules were generated by implanting materials in vivo followed by the harvesting of the capsule for permeability studies, with results of these studies showing lower permeability for thicker capsules [19]. In vivo studies of transplanted pancreatic islets showed that islet viability decreased dramatically with increasing distance between blood vessels and islets [17,20]. This result is not surprising because in a healthy pancreas, islets are very closely associated with capillaries [21]. Increased distance from blood vessels also resulted in lower insulin production from transplanted islets [18]. Devices containing transplanted cells must therefore develop capillary networks with transcapillary mass-transfer rates high enough to ensure survival of the cells. Prevascularization of polymer scaffolds for tissue engineering was attempted as a method to develop a vascular network in and around the implant and promote the survival of cells transplanted into the scaffold at a later date [22]. A number of studies have also shown that the microarchitecture of an implanted material has a great influence on its local tissue response [23]. In suture materials, the topography of the surface affected macrophage involvement, with round sutures resulting in less macrophage involvement than sutures with surface irregularities on the scale of 10–15 µm [24]. In another study, surface features greater than 10 µm appeared to attract foreign-body giant cells to the surface of the implant. The presence of these inflammatory cells around the implanted material resulted in thickening of the fibrous capsule [23]. In addition to implant topography, the anatomical site of implantation influences fibrosis. Acrylic fibers implanted into the subcutaneous tissue of rats showed a higher degree of fibrosis than those implanted into the abdominal fat pad. When islets were introduced, oxygen present inside these hollow fibers was also consumed by islets in minutes [20]. Studies by Padera and Colton [25] and Brauker et al. [26] have shown that certain microporous materials will allow blood vessels to grow and be maintained at the tissue–material interface and in some cases within the pores of the material (Fig. 4). However, this is not true for all porous polymer membranes, even those with similar porosities and chemistries. What appears to be driving the host response is not necessarily the chemistry of the material, but the microstructure of individual features within the material onto which host cells can attach. Materials that are microporous but contain large planar features prompted an avascular host response while the same material lacking these planar features and having a more


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FIG. 4 Top: Micrograph of a tissue section containing a PVDF membrane with a 0.22-µm pore size. Arrows indicate foreign-body giant cells at the tissue material interface. Note the lack of blood vessels near the tissue–material interface. Bottom: Micrograph of a tissue section containing a 5-µm PVDF membrane. Arrows indicate blood vessels either at the tissue–material interface or within the material.

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Recent Advances in Biomaterials fibrous structure prompted neovascularization at the tissue–material interface. If the mechanism of this response can be elucidated, it may then become possible to fabricate materials with a microstructure designed to promote specific levels of neovascularization.

Tissue Engineering Tissue engineering is an interdisciplinary field combining fundamental principles from biology and engineering to understand the structure and function of tissue. Tissue failure or loss due to disease or injury accounts for over $400 billion in total annual health care costs in the United States alone. For some ailments such as burns, kidney failure, and liver failure, tissue transplantation is the conventional treatment. However, there exists a serious shortage in available donor tissue. For example, nearly 40,000 people die each year waiting for a suitable liver for transplantation. To address this need, tissue engineering has developed as a discipline with the goal of restoring, maintaining, and improving the function of living tissue. Tissue engineering has taken on three basic forms. The first is the development of biohybrid artificial organs such as implantable islet-containing bioreactors or microencapsulated islets for the treatment of diabetes. The second is the controlled release of growth factors and cytokines to promote tissue repair or regeneration using existing mechanisms within the body. The third approach, which is addressed here, uses cells from the desired tissue cultured in vitro on a three-dimensional biodegradable scaffold. Once these cells have reached a critical density on the scaffold, they are then transplanted into the subject at the desired location. They will then hopefully continue to grow and mature into the desired tissue, the scaffold having degraded with time. This approach has met with some success, particularly for the generation of skin and cartilage. However, there are several significant obstacles that must be overcome before it can be used for the repair or replacement of many types of tissue. Many of these tissue substitutes take the form of cells from the tissue of interest cultured on biomaterials, frequently microporous biodegradable polymer scaffolds. For example, if one wished to form a tissue-engineered liver, hepatocytes (liver cells) are cultured in vitro on a polymer scaffold to a high cell density. The tissue– material construct is then implanted into the abdomen of the patient. As the tissue continues to grow, the polymer scaffold degrades and is eventually replaced by healthy and functioning liver tissue. Many readers may remember the television news image of a human ear growing on the back of a rat, an example of tissue engineering of cartilage. The ear began as cartilage cells (chondrocytes) cultured on a polymer scaffold in the shape of a human ear. For some types of tissue such as cartilage and skin, tissue-engineered products have moved outside the laboratory and into the marketplace. As an example, tissueengineered skin developed from biodegradable scaffolds are available from Advanced Tissue Sciences. Protein Polymer Technologies is developing silk/elastin gels that can be freeze-dried to form sponges for wound healing and organ-filler applications. Other tissue-engineered products are currently in development, in-


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cluding encapsulated cells and cell-based bioreactors for the treatment of Type I diabetes and liver failure, vascular grafts, and bone replacement. For other types of tissue such as liver and nervous tissue, many challenges are present, including the integration of blood vessels and nerves, adverse tissue responses to the biodegradable polymer, and engineering tissue containing multiple cell types. To promote blood vessel growth in engineered tissue, researchers such as David Mooney of the University of Michigan are developing polymer scaffolds prepared in supercritical carbon dioxide that release blood-vessel-promoting growth factors such as vascular endothelial growth factor (VEGF) [27]. VEGF bioactivity was retained and controlled-release sustained in excess of 30 days. VEGF, also known as vascular permeability factor (VPF), is a protein composed of two identical subunits which has a molecular weight of 45 kDa. This protein is a secreted cytokine that is specific for vascular endothelial cells and has been shown to have angiogenic activity in vivo [28–31]. Numerous studies have also shown that VEGF is secreted by macrophages and that VEGF expression is enhanced in macrophages under hypoxic conditions [31]. Other researchers such as Keith Gooch of the University of Pennsylvania are forming capillaries in vitro that connect to existing blood vessels when implanted. Significant advances have also been made in the engineering of other tissues. For nervous tissue, Schmidt and colleagues found that electrically conducting polymers such as polypyrrole aid the reconnecting of severed nerves [32]. For bone, Anseth and colleagues have developed photo-cross-linkable polyanhydrides that have excellent mechanical properties and can be fabricated into complex shapes such as screws [33]. Oberpenning and colleagues recently tested a tissueengineered bladder (Fig. 5) that, when implanted in dogs, was able to retain urine normally for up to 11 months [34]. This bladder, based on smooth-muscle cells and urethelial cells cultured on a biodegradable poly (lactide-co-glycolide) scaffold, represents the first construction of a hollow organ by tissue engineering. Finally, a number of systems have been developed for the in vivo polymerization of a matrix for tissue engineering, including poly(ethylene glycol) gels that can be interfacially polymerized to prevent intimal thickening following ballon angioplasty [35] and polymer solutions containing chondrocytes that can be polymerized transdermally [36,37].

Drug Delivery Drug delivery research in the controlled release of proteins has become increasingly important as many new drugs are either protein or peptide based. Two prime examples are human growth hormone and erythropoeitin for the treatment of anemia; both of these proteins are billion dollar drugs. Most of these biomolecules are unstable in vivo and as a result must be administered by multiple injections. To stabilize these drugs in vivo and to control the rate at which they are delivered and the location of delivery, new polymers, both degradable and nondegradable, are being developed. These materials are designed to control the mass transfer of drug to the surrounding tissue and to be biocompatible. Many of these materials can maintain a

202


FIG. 5 Radiographic cystograms of a tissue-engineered bladder implanted in a dog. (From Ref. 34.)

constant release of drug over many days and months. Examples of such degradable materials include polyesters such as poly(glycolic acid) used in resorbable sutures, polyanhydrides, and poly(ethylene glycol)–polyester copolymers. These materials are chemically designed such that their degradation products are nontoxic. One example is Gliadel, the first new FDA-approved treatment of gliabastoma (a deadly form of brain cancer) in 20 years [38]. This product is composed of a biodegradable polymer containing the anticancer drug carmustine. After the brain tumor is surgically removed, the space formerly occupied by the tumor is lined with polymer wafers containing the drug. As the polymer degrades over time, the anticancer drug is released directly to the brain in concentrations that cannot be


203

achieved by administering the drug via the bloodstream. As a result, the reoccurrence of the disease is diminished and chemotherapy side effects are reduced because the drug is delivered only where it is needed. Drug delivery products for vaccine delivery, anemia, and cancer treatment are expected to be available shortly. Other research in drug delivery is currently focused on developing targeted polymer delivery systems (such as biodegradable microparticles to the lungs) and developing polymer delivery vehicles for gene therapy. Hydrogels, both degradable and nondegradable, have been of particular interest. These materials are highly swollen with water and can possess equilibrium water content in excess of 90% by weight. With the exception of swelling-controlled or osmotic delivery systems, polymeric materials control drug release by limiting Fickian diffusion of the drug through the solvent-swollen network. This process is highly dependent on polymer chain mobility. Mikos and colleagues recently provided a comprehensive review of both nondegradable and degradable hydrogel systems for drug delivery [39]. Some specific examples are instructive. For example, Khare and Peppas have described the effects of copolymerization, ionic strength, pH, and buffer composition on the release of biomolecules [40]. Particularly in the case of anionic gels, these factors influence the equilibrium water content of the polymer network and hence the mesh size in the gel. Brannon-Peppas and Peppas describe research on the effects of comonomer composition and hydrophobicity on swelling [41]. Network hydrophilicity controls swelling via water absorption, and because equilibrium water content is directly proportional to biomolecule mass-transfer rates out of the gel, extended release profiles can be achieved by increasing the hydrophobicity of the gel. In polyelectrolytic gels, the ionic character is pH dependent and can exhibit substantial changes in swelling caused by shifts in pH [41]. Unless highly localized changes in pH occur, these pH-induced changes in swelling may not occur in vivo. DNA–polycation complexes have received extensive attention in recent years as an alternative to virus-mediated transfection of therapeutic gene therapy vectors into mammalian cells. Polycations investigated to date include polylysine [42], cationic liposomes [43], polyethylene imine [44], and poly(amidoamine) (PAMAM) dendrimers [45]. PAMAM dendrimers are unidispersed and have a high positive-charge density at their surface. Electrostatic complexes of DNA and dendrimers were shown to transfect a variety of cell lines, including fibroblasts, CHO, Rat2, lymphoma, and hepatoblastoma [45]. In the Rat2 cell line, transfection was highly dependent on the generation of the dendrimer, which dictates the surface charge and size of the molecule. Transfection efficiency as compared to simple exogenous DNA was found to increase exponentially with the generation of the dendrimer, reaching a plateau after the ninth generation of 20,000 to 40,000 times more efficient. Although chloroquine and DEAE–dextran enhanced transfection, they were not required and transfection, as determined from the expression of a luciferase gene in the vector, occurred at levels easily measured. Recently, a novel system was described that may have significant applications in drug delivery. Discher and colleagues [46] described the formation of polymersomes, vesicles made of amphiphilic diblock copolymers of poly(ethylene oxide) and poly(ethylethylene) (Fig. 6). These structures were found to be structurally

204


FIG. 6 Polymersomes. (A) Schematic representation of diblock copolymer assembly; (B) micrograph of diblock copolymer vesicles, rodlike structures (black arrow) and micelles (gray arrow). (From Ref. 46.)

tougher than liposomes and significantly less permeable to water. With further modification, these materials may be developed into new vehicles for nucleicacid- and protein-based drugs.

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45. J. Kukowska-Latallo, A. Bielinska, J. Johnson, R. Spindler, D. Tomalia, and J. Baker, ‘‘Efficient Transfer of Genetic Material into Mammalian Cells Using Starburst Polyamidoamine Dendrimers,’’ 93, 4897–4902 (1996). 46. B. Discher, Y. Won, D. Ege, J. Lee, F. Bates, D. Discher, and D. Hammer, ‘‘Polymersomes: Tough Vesicles Made from Diblock Copolymers,’’ Science, 284, 1143– 1146 (1999). MICHAEL V. PISHKO

Recent Development of Extractive Distillation: A Distillation Alternative

Introduction Distillation is the most commonly used method for recovering and purifying petrochemicals and chemicals in the industry. The difference in boiling points between the key components to be separated is the means for separation. The ease of separation is conveniently measured by the relative volatility (α) between the key components, which is defined as: α⫽

Y 1 /X 1 Y 2 /X 2

where X 1 and X 2 are the mole fraction of components 1 and 2, respectively, in the liquid phase and Y 1 and Y 2 are those in the vapor phase. In fact, α is one of the major economic factors for distillation. Colburn and Schoenborn [1] gave the following generalized correlation for the approximate number of theoretical plates required for a separation of products each of 99% ⫹ purity: Number of theoretical plates ⫽

4 α⫺1

In general, distillation becomes uneconomical when 0.95 ⬍ α ⬍ 1.05, because a large number of plates requires very high capital investment and a large reflux ratio requires very high operating cost. Under this situation, a solvent-enhanced distillation, such as extractive distillation (ED), becomes economically attractive and practical. The basis for ED is to increase α by introducing a high-boiling, polar solvent to the distillation column. The characteristics, design, and operation of an extractive distillation column (EDC) have been thoroughly discussed in the literature [2–5].

208

Extractive Distillation Depending on applications, the solvent-to-feed weight ratio (S/F) can vary from 3 to as high as 20, so the EDC is normally operated using substantially higher amount of liquid (nonvolatile solvent) than the conventional distillation. Also, due the to difference in solubility of the key components in the polar solvent, a certain portion of the EDC may have two liquid phases. Significant progress has been made in the industry recently in terms of understanding and handling these special situations in the EDC.

Hydrodynamic Behavior of a Packed EDC The presence of the nonvolatile solvent in the EDC not only substantially increases liquid flow (L) but also reduces the vapor flow (V) by preferentially absorbing the more polar components in the vapor stream. Therefore, extractive distillation is normally operated under significantly higher liquid to vapor ratios (L/V) as compared to conventional distillation. Depending on applications, the solvent-to-feed weight ratio (S/F) varies from 3/ 1 to as high as 20/1. Hydrodynamic behavior of a distillation operation with a highL/V condition has not been significantly reported in the literature. Nevertheless, such information on packed columns used for ED operation was reported by Brown and Lee [6]. Two types of packing were tested; a random dumped packing [0.63-cm protruded metal packing (Pro-Pak)] and a structured packing (Koch-Sulzer BX). Shown in Fig. 1, the tests on Pro-Pak random packings were conducted in a 0.15-

FIG. 1 Schematic diagram of ED pilot plant for hydrodynamic study of packings.

Kettle temp. (°C) 174 174 182 182 182 182 193 193 172 172

Solventto-feed ratio 6.5 6.5 6.5 6.5 10.0 10.0 10.0 10.0 6.5 6.5

Packing code P S P S P S P S P S

ED Runs with Different Tower Packings

4.4 3.2 2.1 3.7 3.6 3.4 1.1 1.6 4.4 3.2

Reflux ratio 99.6 99.1 88.8 99.3 99.2 99.1 72.8 86.7 99.7 99.1

Toluene recovery

99.5 99.8 100.0 100.0 99.9 100.0 100.0 100.0 99.7 100.0

Toluene purity

Extract (%)

95.0 98.3 100.0 100.0 99.3 100.0 100.0 100.0 96.8 100.0

Heptane recovery

95.4 93.3 59.8 95.3 94.4 94.0 46.7 59.3 96.5 93.4

Heptane purity

Raffinate (%)

Notes: Packing codes: P ⫽ protruded packing (ProPak); S ⫽ structured packing (Koch–Sulzer BX). All ratios and percents were measured by weight.

1 1 2 2 3 3 4 4 5 5

Run no.

TABLE 1

Extractive Distillation 209

210

Extractive Distillation m-diameter EDC with 5.7-m total packed height loaded in a duo-column system: The top and bottom halves of the column were connected in series by insulted vapor and liquid lines. During the operation, pipe distributors were installed at the top and at 40% above the bottom of each half-column. The feed (a mixture of 88/12 wt% toluene/heptane) was fed to the location at 55% above the bottom of the entire EDC, and the extractive solvent (N-methyl pyrrolidone) was fed at 88% above the bottom of the EDC. During a normal run, hydrocarbon was introduced into the EDC at a rate of 0.015–0.022 m3 /h at a temperature slightly below the hydrocarbon bubble point and a pressure of 35 kPa. Lean solvent was fed to the top of the upper half of ED tower at a subcooled temperature of 121°C. The solvent-to-feed ratio (S/F) varied from 6.5 to 10 and the kettle temperature varied between 171°C and 193°C. Koch–Sulzer BX structured packing was also tested in the same EDC under the same condition as the Pro-Pak testings. Distributors were placed only at the top of each half-column. A total of 31 elements with 5.3-m total packing height were fitted in the duo-column EDC. A summary of the selected runs is presented in Table 1. To investigate the pressure drop in the EDC under high-L/V condition, a plot of the difference between the measured pressure drop and the predicted pressure drop is given in Fig. 2. It was found that the actual pressure drops measured in this study are significantly higher than the ones predicted by the packing vendor’s correlations [7,8]. The difference between the predicted and actual pressure drop for protruded packing are at least twice that of the structured packing under the same conditions. It was also found that the height equivalent to a theoretical plate (HETP) on both packings was underpredicted by the vendor’s correlations [7,8]. As shown in Fig. 3, the difference in HETP for the structured packing and protruded packing are respectively 5 and 10 in. The actual HETP was roughly estimated by the Fenske– Underwood correlations, utilizing the composition of feed and products, and an average relative volatility. Then, computer simulations were run to determine the number of theoretical stages, using a rigorous computer algorithm capable of simulating distillation processes with multicomponents and multiphases. Experimental activity coefficients

FIG. 2 Comparison of predicted and actual pressure drop.


211

FIG. 3 Comparison of predicted and actual HETP.

for N-methyl pyrrolidone/n-heptane/toluene were used as input to a Renon activity coefficient model, which was used to determine vapor–liquid equilibrium constants. In the computer simulation, an initial trial solution to the computer algorithm was made by assuming constant molal overflow and a guessed (assumed) columntemperature profile. A final solution was then made by convergence on the input specification values, vapor–liquid equilibria, and heat and material balances. A Newton–Raphson convergence technique was used. The gas capacity factor (F s ⫽ the superficial gas velocity times the square root of the gas density) is an important parameter in determining the loading on the column. Figure 4 shows that the pressure drop for both packings increase with increasing F s, but protruded packing consistently shows a pressure drop three times

FIG. 4 Pressure drop versus gas capacity factor.

212


FIG. 5 HETP versus gas capacity factor.

higher than structured packing. The gas capacity factor is also plotted against HETP in Fig. 5. It appears that F s has no effect on HETP for both packings over the range tested. The HETP for protruded packing is slightly higher than that of structured packing (approximately 1–5 in higher). It is estimated from vendor information that the column was at 12% of flood, using Koch–Sulzer BX structured packing [7], and at 32–37% of flood, using the protruded packing [8]. Liquid loading in an EDC is much higher than the conventional distillation column, so it could be the limiting factor in the design of an EDC. The pressure drop is plotted against the liquid loading (kg/h/m2 ) in Fig. 6 for both packings.

FIG. 6 Pressure drop versus liquid loading.


213

FIG. 7 HETP versus liquid loading.

Liquid loading shows a positive linear relationship with pressure drop for the protruded packing and shows a minimum pressure drop at liquid loading around 8070– 8310 kg/h/m2, but otherwise shows little correlation. Liquid loading is also plotted against HETP for both packings in Fig. 7. Although literature sources predict a decrease in HETP with increased liquid loading [9,10], the structured packing had a constant HETP and the protruded packing has little correlation over the liquid loading range investigated. The flow parameter, (L/V)(ρ v /ρ l )0.5, where ρ v and ρ l are the density of vapor and liquid, respectively, is an important parameter when considering the flood point or maximum loading on the column. Assuming the densities (ρ v and ρ l ) to be relatively constant in the range investigated, L/V is plotted against pressure drop and HETP. Figure 8 shows the pressure drop against the maximum L/V predicted in the computer simulation (normally at solvent feed stage). The column packed with the structured packing shows a constant pressure drop over a L/V range of 15–27, whereas the column packed with protruded packing shows a slight decrease in pressure drop with increasing L/V, although the data points were scattered. A very good linear correlation was found when L/V determined by the maximum liquid and vapor loadings in the column is plotted against the pressure drop, as shown in Fig. 9. Both packings show a pressure drop of about 0.1 in. of water per foot of packing for every 0.2 increase in L/V. This is consistent with established principles, as the stage with the maximum L/V plays the most important role in determining pressure drop in a column and will be used for sizing the column. The HETP is also plotted against L/V for the protruded packing in Figs. 10 and 11 for the solvent feed stage and for the stage with maximum vapor and liquid loadings, respectively. Both figures show an increase in HETP with increasing L/V. However, there was no change in HETP with L/V for the structured packing. It is concluded that both random and structured packings show a loss of effi-

214


FIG. 8 Pressure drop versus maximum predicted L/V (on solvent feed tray).

ciency under ED operation (or under very high L/V) as compared to conventional distillation. Both pressure drop per foot of packing and HETP are underpredicted by vendor correlations for both types of packings, although the predicted results for the structured packing is closer to the actual results than those of the protruded packing. Also, compared to random packing, the structured packing appears to be less affected by the changes in gas capacity, liquid loading, and L/V. Further literature data are required to confirm the superiority of the structured packing for extractive distillation services.

FIG. 9 Pressure drop versus L/V determined by maximum liquid and vapor loadings.


215

FIG. 10 HETP versus L/V (on solvent feed tray) for protruded packing.

Handling Two Liquid Phases in EDC One of the considerations in ED technology is the handling of possible formation of two liquid phases in a certain portion of the EDC where the less soluble components are concentrated. The occurrence of a second liquid phase is caused by the fact that some of the less polar components have significantly lower solubility in the polar solvent than the more polar components. One way to solve the problem of two liquid phases in the ED tower is to select

FIG. 11 HETP versus L/V determined by maximum liquid and vapor loadings for protruded packing.

216

Extractive Distillation a polar solvent, which has enough solvency to dissolve all the components in the mixture under process condition. In general, however, solvents with a high selectivity for compounds to be separated will have a reduced solvency (capacity), and vice versa. Therefore, in order to eliminate two liquid phases, one may have to compromise the solvent selectivity, sometimes, to a great extent. A better way is to cope with two liquid phases in the EDC, without sacrificing the solvent selectivity, for following reasons: 1. Although two liquid phases normally reduce the solvent selectivity in a threephase equilibrium (vapor–liquid–liquid) condition in the EDC, it can be compensated by intrinsic selectivity of a highly selective solvent. For example, the performance of sulfolane (SULF) was compared with those of N-formyl morpholine (NFM), N-methyl pyrrolidone (NMP), 2-pyrrolidone (2PD), and dimethyl sulfoxide (DMSO). The rough comparison was made through their abilities to enhance the relative volatility of n-heptane over benzene (an aromatic and nonaromatic separation) in a one-stage equilibrium cell. Table 2 shows that although two liquid phases were observed using sulfolane as the solvent, it still gave a better performance than other solvents when a single liquid phase existed in the mixture. 2. Two liquid phases present no ill effects on the efficiency of small tray or packed towers with diameter from 0.08 to 0.46 m. However, in a larger tower, the heavy liquid phase tends to accumulate on the tray if the liquid phases are not well mixed. This problem can be eliminated by tray designs promoting gas agitation, causing the two liquid phases to behave as a homogeneous liquid that followed general correlations for pressure drop, liquid holdup, broth height, downcomer liquid level, and fractional entrainment. For larger packed

TABLE 2 Comparison of the ED Solvents for n-Heptane and Benzene Separation Phase composition† (wt%) Solvent

S/F

X1

Y1

X2

Y2

α1/2

Liquid phases

SULF SULF NFM NFM NMP NMP DMSO DMSO 2PD 2PD

1.0 3.0 1.0 3.0 1.0 3.0 1.0 3.0 1.0 3.0

16.24 16.32 16.23 16.23 18.60 18.60 15.21 15.21 14.74 14.74

28.94 43.76 26.39 36.84 27.74 35.71 26.97 37.54 24.32 34.76

83.76 83.68 83.77 83.77 81.40 81.40 84.79 84.79 85.26 85.26

71.06 56.24 73.61 63.16 72.26 64.29 73.03 62.46 75.68 65.24

2.10 3.99 1.85 3.01 1.68 2.43 2.06 3.35 1.86 3.08

2 2 1 1 1 1 1 1 1 1

Notes: SULF: sulfolane; NFM: N-formyl morpholine; NMP: N-methyl pyrrolidone; DMSO: dimethyl sulfoxide; 2PD: 2-pyrrolidone. † X1 and X2 are the liquid compositions of n-heptane and benzene, respectively, and Y1 and Y2 are the vapor compositions of n-heptane and benzene, respectively (all in weight% on solvent-free basis). α1/2 ⫽ (Y1 /X1)/(Y2 /X2), the relative volatility of n-heptane over benzene.


217

columns, the liquid–liquid redistributor should be specially designed to allow separate distribution of the two liquid phases [11]. Computer simulations have been developed which are capable of accurately predicting the development of two liquid phases in the EDC and the summary was reported [12]. In one approach, the simulation algorithm starts from linearized pressure, temperature and concentration profiles, and feed conditions given by the program operator. New estimates of composition are solved using the material balance and equilibrium relationship for each tray. Then, the equilibrium constants are reestimated and a new temperature gradient is established to calculate a trayby-tray energy balance. Accumulated errors are calculated for the energy, material, and equilibrium balances. Appropriate column operation restraints are factored in at this point. A correction factor is found for the temperature, rate profiles, and liquid composition profile by inverting the accumulated error matrix. These correction factors are used to form new estimates of composition to start the process again until the correction factors are small enough to call the components converged. Multicomponent vapor–liquid and liquid–liquid equilibria solutions are required for the algorithm. Two activity coefficient models, NRTL and UNIQUAC, are readily extendable to multicomponent systems and capable of such solutions. Experimental activity coefficients, γ, at infinite dilution are used for calculating binary parameters for the NRTL equation. These parameters are then tested using experimental liquid–liquid ternary data, experimental vapor–liquid equilibrium data, and data from pilot plant or commercial plant. The NRTL equation is used in the algorithm to calculate activity coefficients and is given by the following equations:

冤冢

G 21 x 1 ⫹ x 2G 21

冤冢

G 12 (x 2 ⫹ x 1G 12 )

ln γ 1 ⫽ x 22 τ 21

ln γ 2 ⫽ x 21 τ 12

冣

2

⫹

冣

τ 12G 12 (x 2 ⫹ x 1G 12)2

2

⫹

冥

τ 21G 21 (x 1 ⫹ x 2G 21 )2

冥

where ln G 12 ⫽ ⫺β 12 τ 12, τ 12 ⫽

η 12 ⫹ S 12T , RT

ln G 21 ⫽ ⫺β 21 τ 12 τ 21 ⫽

η 21 ⫹ S 21T RT

where G ij , η ij , S ij, τ ij , and β ij are empirical constants, γ i is activity coefficient, R is the gas constant, T is the absolute temperature, and x i is the liquid-phase mole fraction of component i. A Newton–Raphson-based flash algorithm checks for two liquid phases by checking Gibbs free energies for possible second liquid-phase components. If two liquid phases are indeed present, regular solution theory provides a method of combining the liquid-phase activity coefficients.

218


Some Unique Applications of Extractive Distillation Extractive distillation technology has been practiced and continuously improved commercially since World War II when it gained commercial recognition in the recovery of high-purity butadiene, isoprene, and C4 olefins from the C4 and C 5 petroleum streams. However, ED technology for recovering high-purity heavier petrochemicals from the petroleum streams has gained commercial importance only recently. Many ED solvents have been studied during the past 50 years to determine their selectivity for purifying heavier hydrocarbons. As shown in Table 3, a number of solvents selective for aromatic recovery from petroleum streams have been listed in the literature [13]. However, none of the solvents listed in Table 3 have gained commercial importance for BTX aromatic recovery from the petroleum streams. The modern state-of-the-art ED technologies for BTX aromatic recovery are based on several solvent systems: SULF, NFM, and NMP. In most cases, proprietary cosolvents are added to the base solvents to enhance the solvent performance. The modern ED processes can compete very favorably with, for example, liquid–liquid extraction based on sulfolane, which has dominated the BTX aromatic recovery field for many years. However, in the following cases, ED technology may be the preferred or the only choice.

Feeds with High Aromatic Content The recovery of BTX aromatics from pyrolysis gasoline, which contains 80% ⫹ aromatics is one example. The high aromatic content tends to prevent the interface formation between the raffinate and the extract phases in the liquid–liquid extractor, making the process inoperable. Today, two of the leading ED processes for BTX aromatics recovery are offered by GTC Technology Corporation (GT-BTXSM process) and Krupp Koppers (Morphylane process) [14,15]. The following is an example using the GT-BTX process to recover high-purity BTX aromatics from the full-range (C 6 –C 8 ) pyrolyTABLE 3 ED Solvents for Aromatics Recovery from Petroleum Streams Furfural

Acetonyl acetone

Nitrobenzene

Nitrotoluene

Aniline

Dichloroethylether Nonanoic acid

Phenyl cellosolve

Cresol

Sodium-o-xylene sulfonate ⫹ H 2O Propylene glycol o-Chloroamine

o-Phenetidine

Phenol-cresol

Phthalic anhydride Hexyleneglycol o-Chlorophenol Methyl salicylate

2-Ethylhexanol 2-Ethyl hexylamine o-Nitrophenol Dimethyl aniline–aniline

o-Phenyl phenol Dipheyloxide

Phenol


219

sis gasoline, which contains around 90% aromatics. This type of feed is unsuitable for liquid–liquid extraction (LLE), because the aromatic content is so high that it prevents the formation of the interface between the extract and raffinate phases, which is necessary for LLE operation. A part of the raffinate stream from the LLE unit is often recycled to reduce the aromatic content in the feed stream (to ensure the interface formation) and, thus, reduces the efficiency of the process. To illustrate the application of ED technology to high aromatic containing feed, two different feedstocks were tested in a pilot plant consisting of a 60-tray extractive distillation column and a packed solvent stripper for solvent recovery. The simplified feed compositions are given in Table 4. The pyrolysis gasoline feed is introduced to the middle portion of the EDC near its bubble point. Lean solvent is fed near the top of the EDC at about 10°C below the column temperature to generate internal reflux to improve the column performance. The solvent preferentially extracts the more polar components in the mixture, allowing the nonaromatic components to rise as vapor to the top of the column as the raffinate product. The bottoms of the column consist of the solvent and the aromatic components; these are fed to a solvent stripper (containing 9.5 m of random packing) to separate the solvent from the extract products. The lean solvent is then recycled to the top of the EDC. A schematic diagram is presented in Fig. 12. The analyses for the product streams are summarized in Table 5. The solvent-to-feed ratio (weight) for both feeds was 3.0. The recovery for benzene, toluene, and mixed xylenes, were respectively 96.5–97.0%, 99.0%, and 99.9% by weight. In order to determine the purity of BTX aromatics produced without fractionating the extract product, a more detailed analysis was carried out to determine the nonaromatics in the feedstock. Table 6 shows the major components and their boiling points in Feed No. 1. From Table 5, the nonaromatics in the extract (product) stream was 0.54 wt%, and the split of bottom (extract) to feed in the ED column was 0.87. The nonaromatics in the extract is equivalent to 0.47 wt% (0.54 wt% ⫻ 0.87) of the heaviest nonaromatics in the feed. According to Table 6, these nonaromatics were iso-nonanes (0.16 wt%) and the heavy portion of iso-octanes (0.31 wt%). According to material balances, commercial grade of benzene (with 99.9 wt% purity), toluene (with 99.0 wt% purity), and mixed xylenes (with 98.5 wt% purity) can be recovered from the extract product by distillation. The impurities in benzene and toluene products were the 0.31 wt% heavy iso-octanes in the feed, whereas the impurities in mixed xylenes were the 0.16 wt% iso-nonanes in the feed. These experimental results demonstrate the effectiveness of a ED process (GT-BTX process) in recovering chemical grade BTX from a fullrange (C 6 –C8 ) pyrolysis gasoline containing high aromatics. TABLE 4

Composition of Pyrolysis Gasoline for BTX Aromatics Recovery

Component (wt%) Nonaromatics Benzene Toluene C8-aromatics C9⫹-aromatics

Feed No. 1

Feed No. 2

11.50 49.28 27.79 11.35 0.08

7.89 48.83 29.38 13.90 0.00

220


FIG. 12 Schematic diagram of ED process.

Heavy Aromatics Recovery Liquid–liquid extraction is more effective to extract lighter aromatics, such as benzene and toleune, but less effective to extract heavier aromatics, such xylenes and C 9⫹-aromatics. The latter compounds have relatively less solubility in the extractive solvents than the former compounds. Extractive distillation, on the other hand, will provide higher recovery for the heavier aromatics, which tend to stay the solvent at the bottom of the ED column, due to their higher boiling points. However, the solvent boiling point has to be high enough to allow a clean separation between the extract product and the solvent by stripping or distillation in the solvent recovery column. The performance of ED technology for recovering C 8 to C 9⫹ aromatics has not been significantly reported in the literature. To determine the performance of GT-BTX process, two heavy aromatic feeds with composition shown in Table 7 were investigated in a pilot plant consisting

TABLE 5 Composition of Product Streams from the ED Unit for High Aromatic Feeds Component (wt%) Nonaromatics Benzene Toluene C8-Aromatics C9⫹-Aromatics

Feed No. 1

Feed No. 2

Raffinate

Extract

Raffinate

Extract

84.84 13.30 1.87 0.11 0.00

0.54 54.66 31.66 13.04 0.10

80.93 15.69 3.23 0.15 0.00

0.44 52.22 32.06 15.31 0.00


221 TABLE 6

Components and Their Boiling Points of Feed No. 1

Component

Wt%

Boiling point (°C)

Cyclopentane 2-Methylpentane 3-Methylpentane n-Hexane Methyl cyclopentane Benzene Cyclohexane 2-Methylhexane 2,3-Dimethyl pentane 3-Methylhexane 1-cis-3-Dimethyl cyclopentane 1-tr-3-Dimethyl cyclopentane 1-tr-2-Dimethyl cyclopentane n-Heptane 2,2-Dimethylhexane Methyl cyclohexane 2,4-Dimethylhexane Toluene n-Octane Iso-octane Ethylbenzene p-Xylene m-Xylene o-Xylene Iso-nonanes C9-aromatics

0.22 0.55 0.36 1.21 3.57 49.65 1.29 0.30 0.19 0.21 0.20 0.12 0.24 0.32 0.07 0.47 0.52 27.50 0.07 1.36 6.38 0.42 3.32 1.18 0.16 0.38

49.3 60.3 63.3 68.7 71.8 80.1 80.7 90.1 89.8 91.8 90.8 91.7 91.9 98.4 106.8 100.9 109.4 110.6 125.7 107–119 136.2 138.3 139.1 144.4 122–151 152–176

Total

100.26

of a 60-tray extractive distillation column and a packed solvent stripper for solvent recovery. The solvent-to-feed ratio (weight) for both feeds was 3.0. Based on the feed analyses, the kettle temperature and pressure of the ED column were adjusted to achieve the split of bottom-to-feed ratio of approximately 0.87 for Feed No. 1 and 0.76 for Feed No. 2. The solvent stripper (solvent recovery column) was operated

TABLE 7

Composition of Heavy Aromatics Feeds for ED Study

Component (wt%) Nonaromatics Benzene Toluene C8-aromatics C9⫹-aromatics

Feed No. 1

Feed No. 2

12.51 0.31 2.66 63.82 20.71

24.03 0.40 1.52 59.91 14.17

222

Extractive Distillation TABLE 8

Composition of Product Streams from the ED Unit for Heavy Aromatic Feeds Feed No. 1

Feed No. 2

Components (wt%)

Raffinate

Extract

Extract Recovery

Raffinate

Extract

Extract Recovery

Nonaromatics Benzene Toluene C8-aromatics C9⫹-aromatics

85.10 1.95 3.02 8.07 2.15

0.11 0.00 2.59 73.16 23.81

0.75 0.00 83.29 98.06 98.34

87.75 1.51 2.33 8.51 0.05

1.09 0.00 1.31 78.95 19.03

3.34 0.00 63.38 96.91 98.76

with stripping gas under the proper level of vacuum to minimize the stripper temperature. The analyses of the production streams are presented in Table 8. Shown in Table 8, under a solvent-to-feed ratio of 3.0, the overall aromatic recovery and purity have achieved the commercial requirements. The overall aromatic recovery for Feed No. 1 was 97.7 wt%, and for Feed No. 2, it was 95.7 wt%, and the overall aromatic purity for Feed No. 1 was 99.89 wt%, and for Feed No. 2, it was 98.91 wt%. It is also observed that under similar operating conditions, the process performance from Feed No. 1 is indeed better than that of Feed No. 2, probably due to higher aromatic content in Feed No. 1 (87.5 versus 76 wt%).

Revamping Existing LLE Processes Using the ED Method for BTX Recovery The ED method can be most effectively applied to revamping an existing LLE facility for recovering BTX aromatics from reformate or pyrolysis gasoline. This approach has been recently reported by Gentry and Kumar [16]. In a typical LLE process where the solvent has higher density than the feed, the feed is fed into the lower section of the LLE column and flows countercurrently against the solvent. The solvent is introduced to the top of the column and flows downward through the column to preferentially extract the more polar components in the feed into the bottoms. The bottoms stream is then fed to a solvent recovery column to recover the purified more polar components in the overhead and the lean solvent as the bottoms to be recycled to the LLE column. BTX aromatics are recovered using LLE with sulfolane as the extractive solvent. A typical schematic diagram of this type of process is shown in Fig. 13. Reformate containing BTX aromatics and the extractive solvent are fed into the LLE column according to the manner described in previous paragraph. The solvent extracts the aromatics and some nonaromatics into the bottoms, which is fed to the extractive stripping column. This column strips the nonaromatics from the solvent and aromatics, for recycle to the LLE column. The bottoms from the extractive stripping column contain the aromatics and solvent are separated in a solvent recovery column. Lean solvent from the solvent recovery column is recycled back to the LLE column. The raffinate stream from the top of LLE column contains some solvent, which is recovered in a water-wash column. The water is returned to the system in a closed-loop recycle.


223

FIG. 13

Schematic diagram of LLE process using sulfolane solvent.

A fundamental aspect of this process is that the solvent exhibits a selectivity favoring lower-boiling components more than the high-boiling components. The solvent selectivity favors the hydrocarbon species according to the following sequence: aromatics ⬎ naphthenes/olefins ⬎ paraffins. Lighter nonaromatic impurities are more likely to be coextracted due to their solvent affinity in LLE column, but they should be the easiest to remove in the ED column due to their lower boiling point. Adding just one ED column to the existing LLE process system can not only substantially increase the process throughput but also improve the performance in terms of product quality and purity. The following are some examples.

Case A: Purifying By-product Benzene The xylene isomerization and toluene disproportionation units within the aromatic complex produce benzene as a by-product, but the quality of benzene is low and requires reprocessing in the LLE unit to upgrade its purity. Figure 14 shows a new approach using a hybrid of the LLE process with ED that bypasses part of the feed around the original extraction section.

224


FIG. 14 A hybrid of LLE and ED process purifying by-product benzene.

In the hybrid scheme, the by-product (benzene-rich feedstock) is fed to the stand-alone ED column, where the maximum aromatics limit in the feed charge is not a concern. The original feed to the extraction section is not affected by this new ED operation. The rich solvent from both LLE and ED operations are combined into the existing solvent-recovery column. The typical solvent-recovery column can usually manage the higher capacity or can be redesigned to do so. If desired, the raffinate from the two can be segregated for optimum disposition; the ED raffinate rich in cyclohexane may be recycled to the reformer unit, whereas the LLE raffinate rich in paraffins could be blended into the feedstock for a naphtha cracker. The conventional LLE process can be retrofitted to use this hybrid process, without requiring extensive modifications, investment, or shutdown time. The primary changes are modifying the solvent system to be compatible with both LLE and ED operations and to make the appropriate tie-ins to the ED column.

Case B: Optimize the BTX Product Distribution Among the BTX aromatics, toluene has historically contributed the least upgrade over its alternative value in motor fuel (e.g., the petrochemical value of toluene


225

FIG. 15 A hybrid of LLE and ED process for selective purge of toluene.

is only marginally above its fuel blending value); thus, it is not profitable to extract toluene. In a conventional LLE unit for BTX aromatics recovery, toluene is inevitably extracted along with benzene and xylenes. Figure 15 shows a LLE and ED hybrid scheme where toluene is selectively purged from a BTX mixture to shift the product mix toward benzene and xylenes and to avoid some operating cost for toluene. As shown in Fig. 15, the reformate splitter is retrofitted to include a side cut of the C 7 and C 8 components, which is fed to an ED column operated in parallel with the existing LLE unit. Operation of the ED column is made to intentionally purge the majority of the toluene along with the nonaromatics to the column overhead. Then, the toluene-lean ED column bottoms are combined with the benzene-rich solvent mixture from the extractive stripper bottoms as feed to the solvent-recovery column. A number of benefits from this operation can be realized: (1) overall processing costs are reduced because the stream contains less toluene; (2) xylenes recovered should have higher purity because the majority of the C 8 –C 9 nonaromatics are purged with toluene in the ED column; (3) substantial increase in overall process capacity by adding only one ED column to the existing LLE system.

Case C: Capacity Increases Adding a new ED column to an existing LLE unit (as shown in Fig. 13) can double the capacity of the unit. As demonstrated in Fig. 16, the new ED column takes

226


FIG. 16 A hybrid of LLE and ED for substantial increase of throughput.

the fresh feed while the main liquid–liquid extractor takes the feed from the ED column overhead raffinate stream. Solvent is divided between the ED column and the raffinate extractor. The original extractive stripper column is converted into a solvent-recovery column operating in parallel with the existing solvent-recovery column. Using a single new ED column to retrofit the existing LLE unit can create significant increase in throughput and improvement in product quality with minimized capital costs.

Styrene Recovery from Pyrolysis Gasoline One of the most difficult separations, using the ED method, is the purification of styrene from the close-boiling C 8 aromatic isomers. The method is based on the slightly higher polarity in styrene than the other C 8 aromatics, due to the double bond in the side chain of the styrene molecule. Recently, a proprietary ED process was developed for the commercial application for recovering styrene directly from pyrolysis gasoline [17]. In developing this process, a number of extractive solvent candidates were screened in the laboratory for selectivity, solubility, and other important properties, such as thermal stability, toxicity, corrosivity, boiling point, freezing point, and so forth. The selected solvent underwent an extensive pilot-plant test program to determine its performance in the ED process, and, consequently, to optimize the


227 TABLE 9 Composition of Feed and Products of the ED Pilot Plant for Styrene Recovery Component Dicyclopentadiene Benzene Toluene Vinylnorbornenes Ethyl Benzene p-Xylene m-Xylene Cumene o-Xylene Styrene Allylbenzene Nonaromatics 2,5-Dimethylthiophene Solvent

Feed (wt%)

Overhead (wt%)

Bottom (wt%)

0.26 0.22 35.81 3.73 12.03 2.97 7.45 395 ppm 4.52 23.82 390 ppm 9.08 172 ppm 0

0.40 0.23 43.68 4.25 13.47 3.30 8.80 431 ppm 5.79 10.31 269 ppm 9.72 195 ppm 0.37

⬍1 ppm ⬍1 ppm ⬍1 ppm ⬍1 ppm ⬍1 ppm ⬍1 ppm ⬍1 ppm ⬍1 ppm ⬍1 ppm 1.11 23 ppm ⬍1 ppm ⬍1 ppm 98.88

key process variables to support the computer process simulation for scale-up and commercial design. The pilot-plant tests were conducted in a 7.6-cm-diameter ED column packed with 7.3 m of knitted wire-mesh packing (Goodloe Style #773). The C 8 cut of pyrolysis gasoline with the composition shown in Table 9 was fed at a location 4.3 m from the column bottom and the lean solvent was introduced at a location 6.4 m from the column bottom. A schematic diagram of the pilot plant can be represented by Fig. 12. The column was operated under reduced pressure to minimize the column operating temperature, and a proprietary inhibitor was added in trace amounts to the feed to prevent styrene polymerization in the column. The key variables, such as S/F, the reflux ratio, and the kettle temperature, were properly adjusted to yield styrene with 99.9 wt% purity (on a solvent-free basis) in the column bottoms. The composition of the column overhead and the bottom products are also given in Table 9. In this operation, a cosolvent was used as the key to improve the styrene purity from 95 to 99.9 wt%, to reduce the column temperature by 35°C, and to decrease the solvent circulation by 20%. The bottom product (the rich solvent) was routed to the solvent stripper column, where styrene was distilled overhead, and the lean solvent exited from the bottom of the column for recycle to the ED column. The separation strategy for recovering styrene from pyrolysis gasoline is to remove the majority of the light and heavy components in pyrolysis gasoline by conventional distillation and to use ED to remove the remaining close-boiling components. Figure 17 shows a possible scheme for processing pyrolysis gasoline, which includes styrene recovery by ED. The main benefits for recovering styrene from pyrolysis gasoline generated from a naphtha cracking complex are as follows:

228


FIG. 17 Schematic diagram for pyrolysis gasoline processing including styrene recovery.

1. The styrene component is upgraded from motor fuel value to petrochemical value. 2. The xylenes can be upgraded from motor fuel value to the feedstock value to xylene isomer unit. 3. The overall hydrogen consumption to convert styrene to ethylbenzene is reduced. 4. The catalyst fouling and operating costs in the selective hydrotreater are reduced. 5. The potential debottleneck of the hydrotreating area can be avoided; it otherwise would be required for naphtha cracking expansion. The economics of styrene recovery from pyrolysis gasoline depend to a great extent on the quantity of styrene available in this stream. Units produce at 10,000 to 15,000 metric tons per year of styrene in pyrolysis gasoline would be the candidate for this technology. Table 10 shows that the economics for styrene recovery with a pretax return on investment (ROI) of 41% can be achieved, based on a 25,000-metric tons per year plant capacity.

Cyclohexane Recovery from Natural Gas Liquid or Naphtha Cyclohexane exists naturally in naphtha and natural gas liquid (NGL) streams and is an important raw material for the nylon industry. As shown in Table 11, it is impossible to recover high-purity cyclohexane from these streams by conventional distillation, because of the close-boiling C 7 iso-paraffins in the streams. Because


229 TABLE 10

Economics for Styrene Recovery from Pyrolysis Gasoline

Typical U.S. Gulf Coast capital cost Styrene value in pyrolysis gasoline Styrene product sales value Processing cost Gross margin Pretax return on investment

$ 20 million $ 180 per metric ton $ 550 per metric ton $ 40 per metric ton $ 8.3 million per year 41%

Note: Basis: 25,000 metric tons per year styrene capacity from a world-scale naphtha cracker. Values on styrene product, feedstock, processing cost, and capital investment were calculated based on the 1997 published information.

cyclohexane and the close-boiling components in the feed mixture have only a very small difference in polarity and are relatively insoluble in the selective solvents, it also takes a very difficult ED process to do the separation. Because no effective single ED solvent has been found, a mixed solvent was developed commercially to recover high-purity cyclohexane directly from an NGL fraction containing 85% cyclohexane [18]. The proprietary mixed solvent (the MIST solvent) was developed through extensive test in the laboratory and the evaluation in a 150-mm-diameter ED pilot plant using a refinery NGL stream, which has an average composition shown in Table 11. In fact, the cosolvent in MIST solvent played a major role in the success of this process [18]. As demonstrated in Fig. 18, The overall cyclohexane recovery changed from 100 to 56 wt% as the cosolvent concentration decreased from 30 to 10 wt% under a constant kettle temperature of the ED column. Meanwhile, the recovery of 2,4-dimethylpentane in the raffinate stream (ED column overhead) increased from 87 to 96.3 wt% over the same composition range of the MIST solvent. The higher 2,4-dimethylpentane in the raffinate stream, the higher cyclohexane purity in the extract (product) stream. Table 12 summarizes the results for 80% and 90% cyclohexane recoveries at various cosolvent concentrations. At an 80% cyclohexane recovery, MIST solvent TABLE 11

Average Composition and Boiling Point of the Feedstock for Pilot-Plant Testing

Component Cyclohexane 2,2-Dimethylpentane 2,4-Dimethylpentane 3,3-Dimethylpentane 2,3-Dimethylpentane 2-Methylhexane 3-Methylhexane 2,2,3-Trimethylbutane Dimethylcyclopentane n-Heptane

Weight%

Boiling point (°C)

89.1 1.3 4.0 0.1 0.9 1.6 1.1 0.8 1.0 0.1

80.7 79.1 80.4 86.0 89.7 89.9 91.9 80.8 90.6 98.3

230


FIG. 18 Effect of cosolvent in MIST solvent on product purity and recovery.

containing 10% cosolvent appears to give the greatest cyclohexane purity. At 90% cyclohexane recovery, MIST solvent having 25% cosolvent had similar performance to that with 10% cosolvent. Finally, the MIST solvent with 25% cosolvent showed better performance than the one with 10% cosolvent. Based on the successful pilot-plant testing on the MIST solvent, a commercial plant for purifying 100 metric tons per day cyclohexane was designed, constructed and started up in 1991. TABLE 12 Effect of Cosolvent Concentration on Cyclohexane Recovery and Purity

Cosolvent (wt%) 25 20 10 30 25 10 25 10

Cyclohexane Recovery (wt%)

Cyclohexane Purity (wt%)

2,4Dimethylpentane Recovery (wt%)

80 80 80 90 90 90 94 94

98.9 99.0 99.1 99.2 99.2 99.2 99.1 99.1

89.8 92.0 93.0 87.2 90.0 90.3 86.8 80.0

Polyolefins Produced by Single-Site Catalyst

231

References

1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13. 14. 15. 16. 17. 18.

A. P. Colburn and E. M. Schoenborn, Trans. AIChE, 43, 42 (1945). G. T. Atkins and C. M. Boyer, Chem. Eng. Prog., 45, 553 (1949). J. M. Chambers, Chem. Eng. Prog., 47, 555 (1951). K. H. Hackmuth, Chem. Eng. Prog., 48, 617 (1952). R. M. Butler and J. A. Bichard, U. S. Patent 3,114,783 (1963). R. E. Brown and F. M. Lee, ‘‘Effect of Packing on Distillation Columns with High Liquid to Vapor Ratios,’’ AIChE Annual Meeting, Miami Beach, FL, 1992. R. Billet, ‘‘Packed Column Analysis and Design,’’ Department for Thermal Separation Processes, Ruhr-University Bochum, 1989, p. 11. Cannon Instrument Co., ‘‘Pro-Pak Protruded Metal Distillation Packing,’’ Bulletin 23, Cannon Instrument Co., State College, PA. R. Billet, ‘‘Packed Column Analysis and Design. Department for Thermal Separation Processes,’’ Ruhr-University Bochum, 1989, p. 41. D. P. Kurtz, et al., Chem. Eng. Prog., 87,(2), 43 (1991). C. C. Herron, Jr., B. K. Kruelskie, and J. R. Fair, AIChE J., 34,(8), 1267 (1988). B. A. Todd and F. M. Lee, ‘‘Two Liquid Phases in Extractive Distillation for Aromatic Recovery,’’ AIChE Summer National Meeting, Seattle, WA, 1992. M. Van Winkle, Distillation, McGraw-Hill, New York, 1967, p. 464. J. C. Gentry and F. M Lee, ‘‘New Extractive Distillation Process for Aromatics Recovery,’’ AIChE Spring National Meeting, Houston, TX, 1995. G. Emmrich, ‘‘Morphlane: Operational Experience and Results Obtained with New Plants,’’ NPRA Spring Meeting, San Francisco, 1995. J. C. Gentry and C. S. Kumar, Hydrocarbon Process, 69, (March 1998). F. M. Lee and J. C. Gentry, Hydrocarbon Eng., 3(6), 62 (1998). R. E. Brown and F. M. Lee, Hydrocarbon Process, 83 (May 1991).

FU-MING LEE

Structure, Properties, and Applications of Polyolefins Produced by Single-Site Catalyst Technology

Introduction Polyethylene is composed of only carbon and hydrogen (with some exceptions), which can be combined in number of various ways to make many different polyethylenes. These can generally be grouped into six types:

232


• • • • • •

LDPE, low-density polyethylene EVA, ethylene vinyl acetate copolymers HDPE, high-density polyethylene LLDPE, linear low-density polyethylene ULDPE, ultralow-density polyethylene (also known as Very low-density polyethylene, VLDPE) Single-site polyethylenes (e.g., substantially linear homogeneous polyethylene)

Polyethylene technology encompasses a range of crystallinity and melting points. Figure 1 illustrates the approximate range of densities, crystallinities, and melting points available within each major type of polyethylene. Specialty ethylene copolymers such as acid copolymers and ionomers are not included. The total polyethylene consumption is expected to exceed 100 billion pounds by the year 2000. Applications include films, extrusion coatings, injection-molded and blowmolded parts, fibers, adhesives, wire and cable coatings, large parts such as storage tanks made by rotational molding process, foams, and so forth. The most recent advance in polyethylene technology is the development and commercialization of polyethylenes produced by single-site-catalyst technology. Two major classes of single-site catalyst (SSC) technology developed for the polymerization of ethylene and α-olefins are the metallocene catalyst (MTC) and the constrained-geometry catalyst (CGC) systems. The use of these catalyst technologies has allowed a very rapid development of olefin copolymers with a wide range of structures and related properties. This technology has initiated a major revolution for the polyolefin industry [1–4]. Several families of MTC- and SSCtechnology-based polyolefin copolymers have been commercialized in the 1990s. These include polyolefin elastomers (e.g., ENGAGE from DuPont Dow Elastomers, LLC), polyolefin plastomers (e.g., AFFINITY from The Dow Chemical Company; EXACT from Exxon Chemical Company), EPDM (NORDEL IP from DuPont Dow Elastomers, LLC), enhanced polyethylene (ELITE from The Dow Chemical Company); gas-phase LLDPE (EXCEED from Exxon Chemical Company); slurry LLDPE (mPACT from Phillips Petroleum Company), and

FIG. 1 The approximate range of density, crystallinity, and melting point available within each major type of polyethylene.


233

polypropylenes (ACHIEVE from Exxon Chemical Company). In addition to these commercial activities, several other single-site-catalyst-related technologies that allow the copolymerization of α-olefins with polar comonomers are also under industrial and academic development [5,6]. This article will focus on the solid-state structure, rheology, properties, and typical applications of homogeneous polyolefin copolymers made by copolymerization of ethylene and α-olefins (e.g., 1-butene, 1-hexene, 1-octene) using SSC technology. Unlike conventional linear low-density polyethylenes (LLDPE) made by copolymerization of ethylene and α-olefins with Ziegler–Natta (Z-N) catalysts, ethylene–α-olefin copolymers produced by SSC technologies have narrow composition distributions (narrow molecular-weight and comonomer distributions). Hence, they are called homogeneous copolymers and they behave much more like ideal polymers. The polymerization kinetics and the resulting polymer and copolymer structures can be modeled. This significantly advances the fundamental understanding of the structure–property relationships. The major advantage of singlesite catalysts is its versatility in building well-defined molecular structures. This capability allows the polymer and material scientists to design new polymers using a molecular architecture approach and to develop new products with exceptional speed.

Preparation of Ethylene–α-Olefin Copolymers by SSC Technology Two major families of high-efficiency SSC are commercially used for the preparation of polyethylene copolymers. These are a bis-cyclopentadienyl (Bis-Cp) single-site metallocene catalyst (also known as a Kaminsky catalyst) and a halfsandwich, constrained-geometry mono-cyclopentadienyl single-site catalyst (known as constrained-geometry catalyst, CGC, under the trademark of INSITE by The Dow Chemical Company). These two catalyst systems are illustrated in Fig. 2. The Zr in the Zr-based MTC used by many polymer producers has an

FIG. 2 Structure of the Bis-Cp metallocene catalyst and the constrained geometry catalyst.

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Polyolefins Produced by Single-Site Catalyst oxidation state of ⫹4 and the Ti in the Ti-based CGC system has an oxidation state of ⫹2. Single-site-catalyst technology polyolefin copolymers can be produced by high pressure, solution, gas-phase, and slurry polymerization processes. Typical process conditions for making polyolefin copolymers in these processes are as follows: 1. High-Pressure Process: Two types of commercial reactor, stirred autoclave and a tube reactor, are used for the polymerization of ethylene and ethylene– α-olefin copolymers at high pressure. Ethylene and an α-olefin comonomer are usually compressed to at least 10,000 psi when fed into the reactor. The polymerization temperature usually exceeds 100°C. 2. Solution Process: Stirred reactors are usually used for the polymerization of ethylene and ethylene–α-olefin copolymers in a solution phase. In most cases, C6 –C8 hydrocarbons are used as the solvent. The reactors are operated under about 500 psi pressure and the polymerization is carried out at ⬎60°C. 3. Gas-Phase Process: In the gas-phase process, ethylene and α-olefin comonomers (1-butene and/or 1-hexene) are polymerized in the solid state in a fluidized-bed reactor into a powder form. The polymer powder is then converted into pellet form by an extrusion process. The reactor pressure is usually set at ⬎300 psi and the polymerization process is generally carried out at ⬍90°C. 4. Slurry Process: In the slurry process, polymers are made in stirred or loop reactors with an organic liquid carrier (C 4 –C6 hydrocarbons). The polymerization process temperature is generally carried out at ⬍90°C at a reactor pressure of ⬍300 psi. Products made in this process are in powder form and are converted into pellets using extrusion processes.

Molecular Structure of Polyethylene Copolymers Made by Single-Site-Catalyst Technology Molecular-Weight Distribution Ethylene–α-olefin copolymers prepared by single-site-catalyst technologies exhibit a narrow molecular-weight distribution (MWD). MWD, or polydispersity, is the ratio Mw /Mn, where Mw is the weight-average molecular weight and Mn is the number-average molecular weight. Copolymers prepared by a conventional Zeigler–Natta catalyst have a MWD usually larger than 3, whereas the MWD of homogeneous polyolefins made by SSC technology is usually less than 2.5. Figure 3 compares gel permeation chromatography (GPC) traces of a conventional, heterogeneous ethylene–octene copolymer versus a homogeneous copolymer. Polymers with a narrow MWD in general have increased toughness and less ‘‘solvent extractables.’’ On the other hand, the narrow MWD of a linear homogeneous co-


235

FIG. 3 GPC MWD comparison of two ethylene–octene copolymers having a 0.920-g/cm3 density.

polymer also results in poor melt processability (low melt strength, high extruder back-pressure, high-energy consumption during extrusion, etc.). Crystallinity The comonomer content in the copolymer has a profound effect on the properties of the polymer, including crystallinity, thermal, and mechanical properties. In the case of ethylene–α-olefin copolymers, the crystallinity is measured and specified using density. The density of the amorphous phase (ρa) and crystalline phase (ρc) of ethylene–α-olefin copolymers at room temperature is about 0.853 g/cm3 and 1.000 g/cm3, respectively. Weight percent crystallinity can be obtained from the measured density (ρ) as follows: Wt% cryst. ⫽

冢

冣

ρc ρ ⫺ ρa ρ ρc ⫺ ρa

(1)

The effect of comonomer content on the density of homogeneous ethylene–octene and ethylene–butene copolymers, made using a CGC catalyst, is illustrated in Fig. 4. The comonomer content was measured using nuclear magnetic resonance (NMR). Higher mole% butene is required to achieve the same density compared to octene, especially at lower densities. Comonomer Distribution Because of the significant effect of the α-olefin comonomer on the properties of the polymer, it is critical to understand and to measure the comonomer distribution

236


FIG. 4 Densities of homogeneous ethylene–octene and ethylene–butene copolymers as a function of mole% comonomer.

among the polymer molecules (intermolecular distribution) and along the backbone (intramolecular distribution). Intermolecular comonomer distributions can be measured by the temperature rising elution fractionation (TREF) technique [7]. In this procedure, a heated solution of the polymer is placed in a column, packed with very small stainless-steel shot, and slowly cooled to room temperature. The temperature is then increased while solvent flows through the column to a detector that records the amount of polymer in the solvent at any given temperature. As illustrated in Fig. 5, the intermolecular comonomer distribution for a homogeneous ethylene–octene copolymer prepared by the CGC technology is much narrower than that of a conventional LLDPE, which is a mixture of polymer molecules having different levels of an α-olefin comonomer incorporated in the polymer backbone. The narrow TREF curve for the homogeneous copolymer signifies that the number of comonomer units per unit chain length between the copolymer molecules is very similar. However, although the intermolecular comonomer distribution for the homogeneous polymer is very narrow, this does not mean that the intramolecular comonomer distribution is uniform. The narrow intermolecular comonomer distribution for the homogeneous polymer arises from the single-site nature of the MTC and CGC. The uniformity of the intramolecular comonomer distribution, however, is dictated by the reactivity ratio of the monomer and the comonomer. Intramolecular comonomer distribution in homogeneous copolymers made by SSC technology can also have an effect on the solid-state structure and properties of the polymer (e.g., thermal properties, optics, etc.). To address the issue of the effect of the intramolecular comonomer distribution on polymer properties of a homogeneous polymer, structural characteristics for


237

FIG. 5 Intermolecular comonomer distribution of ethylene–octene copolymers measured by TREF. (From Ref. 10.)

CGC ethylene–octene copolymers were modeled using the Monte Carlo simulation [8]. Figure 6 illustrates the intramolecular comonomer distribution for three CGC copolymers, from 0.87 g/cm3 to 0.92 g/cm3 density, in terms of ethylene block length between the hexyl branches within a polymer molecule. As illustrated in Fig. 6, the intramolecular comonomer distribution for the higher-density polymer (0.92 g/cm3), in terms of ethylene block-length distribution between the short branches that are formed from the α-olefin comonomers, is much broader than that of the lower-density copolymers. This is due to the fact that the lower-density

FIG. 6 Model prediction for intramolecular comonomer distribution of three CGC ethylene–octene copolymers: ethylene block-length distribution.

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Polyolefins Produced by Single-Site Catalyst copolymers have more comonomer units along the polymer backbone. This results in a shorter ethylene block length between the short branches, and the distribution is also narrower.

Long-Chain Branching Homogeneous ethylene homopolymers and ethylene–α-olefin copolymers first made by CGC technology have one more unique molecular structural feature not found in the original MTC technology polymers. This unique molecular structural feature is long-chain branching (LCB). Homopolymers and copolymers made by the CGC technology at certain process conditions contain a controlled amount of long-chain branching. These polymers are referred to as substantially linear homogeneous polyethylenes. To summarize this, the molecular structure of three ethylene–α-olefin copolymers (conventional LLDPE made by a Z-N catalyst, homogeneous copolymers made by MTC, and substantially linear homogeneous copolymers made by CGC) are schematically illustrated in Fig. 7. One possible mechanism of LCB formation via CGC technology is as follows. Thermal termination is one of the most effective ways to control molecular weight during the polymerization process. In the thermal termination process, polymeric molecules with a vinyl chain end are formed due to a ‘‘β-hydride elimination’’ mechanism that is taking place in the thermal termination step (Fig. 8). Depending on the last monomer or comonomer insertion, the unsaturation could be vinyl, vinylidene, or trans as shown in Fig. 8. The polymer with a vinyl chain end becomes one of the reactants (in addition to ethylene and α-olefin) that reacts with the catalyst site and forms long-chain branching. A polymer with LCBs made by the CGC technology has many unique rheological properties [9,10]. A few of the most significant features for the CGC-technology polymer are its melt fracture resistance and the control of shear-thinning behavior, as illustrated in Figs. 9 and 10 respectively. In Fig. 10, I2 is the melt index (MI) of the polymer (flow rate in grams per 10 min at 2.16 kg weight, at 190°C) and

FIG. 7 Molecular structure comparison of three different ethylene–α-olefin copolymers. (EXACT is a trademark of Exxon Chemical Company. AFFINITY is a trademark of The Dow Chemical Company.)


239

FIG. 8 Reactor sources of unsaturation in polyethylene during polymerization of ethylene and αolefin comonomers. (From Ref. 10.)

I10 is the flow rate in grams per 10 min at 10 kg weight, at 190°C. The I10 /I2 ratio is a measure of shear-thinning behavior of the polymer melt.

Solid-State Structure and Morphology of SSC-Technology Ethylene Homopolymers and Ethylene–α-Olefin Copolymers The ethylene homopolymer and ethylene–α-olefin copolymers are semicrystalline polymers. Three major factors that affect the crystal morphology and crystallinity of ethylene polymers in the solid state are (1) the amount and size of α-olefin in the polymer, (2) the molecular weight of the polymer, and (3) the crystallization

FIG. 9 Extrudates of CGC polymer and LLDPE at 3.66 ⫻ 106 dyn/cm2 shear stress. (Irgafos is a trademark of Ciba-Geigy Corporation.)

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FIG. 10 Rheology of CGC polymers measured at 190°C: effect of LCB on shear-thinning behavior of three 1 MI polymers with various amount of LCB. LCB/10,000°C was estimated using a kinetic model.

conditions (crystallization temperature at isothermal condition and/or cooling rate at nonisothermal crystallization conditions) [11]. A recent review article on polymer crystals was written by Phillips [12]. The most generally accepted crystal morphology for ethylene polymers is lamellae formed from a folded-chain conformation [13–15]. For very low-density SSC-technology ethylene–α-olefin copolymers (which contain a high percentage of comonomer), the crystal morphology can be very different from the conventional lamella model. For example, let us consider the morphological model of a 60/40 wt% ethylene–octene copolymer made by CGC technology. At this level of comonomer content, the polymer exhibits about 10% crystallinity, as measured by differential scanning calorimetry (DSC) and a density of approximately 0.87 g/cm3. The intramolecular comonomer sequence distribution of the hexyl branches from the octene comonomer, calculated from the reactive-ratio kinetic model (Fig. 6), illustrates that more than 90% of the octene units are less than 50 ethylene units apart. The regular ethylene block length between the octene units is, therefore, much less than the minimum length to form one ˚ thick). For folded unit to form the thinnest possible lamella (which is about 30 A a chain with this molecular structure, it is expected that the polymer has to crystallize in a crystalline form different from the conventional folded-chain lamella model. Figure 11 is a transmission electron micrograph of the crystal morphology of a higher-density (0.920 g/cm3) ethylene–octene copolymer made by CGC technology. At this density and octene level, the copolymer clearly shows a lamellar morphology. Figure 12 is a transmission electron micrograph of a very low-density ethylene–octene copolymer (0.87 g/cm3). It clearly shows that the major crystalline structure of this very low-density, low-crystallinity copolymer is ‘‘spotlike’’ fringed micelle crystals. Polymers having higher crystallinities with the lamella-


241

FIG. 11 Transmission electron micrograph of a 0.920-g/cm3 ethylene–octene copolymer made by CGC technology.

FIG. 12

Transmission electron micrograph of a 0.87-g/cm3 ethylene–octene copolymer made by CGC technology.

242


FIG. 13 Classification of SSC ethylene–α-olefin copolymer based on crystal morphology. (From Ref. 16.)

type crystal morphology are expected to have higher moduli and undergo yielding when deformed. Polymers having low crystallinities with the fringed micelle crystals, however, are expected to be more like elastomers and do not have well-defined yielding behavior. With these two different types of crystal structure, the SSCtechnology homogeneous ethylene–α-olefin copolymers can, therefore, be classified into four major domains [16], as illustrated in Fig. 13.

Thermal and Dynamic Mechanical Properties Melting Behavior of SSC Polymers Differential scanning calorimetry is one of the most popular methods for studying the thermal properties (melting and crystallization) of semicrystalline polymers. Structural information can be uncovered by a careful interpretation of DSC thermograms generated at different conditions. Due to the narrow intermolecular comonomer distribution, SSC copolymers usually have a much narrower melting peak than their heterogeneous counterparts produced by multiple-site Z-N catalysts. Figure 14 shows the DSC melting curve of two ethylene–octene copolymers, both at similar density: one made by a conventional multiple-site Z-N catalyst and one by SSC technology. The heterogeneous LLDPE copolymer always has a peak melting point at around 120–130°C over a broad density range (⬃0.89–0.95 g/cm3). This is because the conventional heterogeneous LLDPE copolymer is a reactor blend of molecules ranging from the un-


FIG. 14

243

DSC melting curve of SSC and LLDPE ethylene–octene copolymers at about 0.903 g/cm3 density. Samples were cooled and heated at 10°C/min.

crystallizable and poorly crystallizable type to the HDPE type containing only of α-olefin [17]. These HDPE-type linear chains can, therefore, crystallize to a rather large size crystal that has a melting point in that temperature range. The shape of the melting peaks of all the homogenous copolymers still represents a single peak or a single peak with a shoulder, very different from that of the heterogeneous LLDPEs, as illustrated in Fig. 15. The peak melting point of the homogeneous copolymer, however, drops accordingly as the density/crystallinity of the polymer is lowered, as seen in Fig. 15. This characteristic has a very significant commercial value for heat-seal and oriented-shrink-film food-packaging applications. A plot of the melting points of LLDPE/VLDPE, LDPE, SSC ethylene–octene, and EVA resins as a function of density is shown in Fig. 16. All the samples were cooled and heated at 10°C/min using a Perkin-Elmer DSC-7. LDPE resins have slightly lower melting peak temperature compared to SSC ethylene–octene copolymers made using CGC technology at a given density. The density of EVA resins increases as the vinyl acetate content increases, even though the degree of crystallinity decreases. This is due to the bulky nature of the vinyl acetate group increasing the amorphous phase density. Hence, the density of EVA resins cannot be compared with the density of ethylene–α-olefin copolymer resins. A plot of melting points of LDPE, LLDPE/VLDPE, and SSC ethylene–octene and EVA resins as a function of resin crystallinity is shown in Fig. 17. Prior to 1990, EVAs were the only cost-effective polyethylenes widely available for packaging applications requiring low-temperature sealing, low-temperature shrinkage, and toughness properties. This is due to their lower crystallinity (typically less than about 40%) and lower melting point (typically less than about 100°C). Beginning in

244


FIG. 15 DSC melting curve of SSC and LLDPE ethylene–octene copolymers at various densities. Samples were cooled and heated at 10°C/min.

FIG. 16 Peak melting temperatures of heterogeneous and homogeneous ethylene–octene copolymers, LDPE, and EVA resins versus polymer density.


245

FIG. 17 Peak melting temperatures of heterogeneous and homogeneous ethylene–octene copolymers, LDPE, and EVA resins versus polymer crystallinity by DSC.

early 1990, polyolefin plastomers (POPs) and polyolefin elastomers (POEs), made using a single-site catalyst, have become available over a wide range of crystallinity and melting points. When plotted against resin crystallinity, the melting points of EVA resins lie very close to the curve for the SSC ethylene–octene resins (Fig. 17). Hence, SSC resins can potentially be used in various film applications where EVAs are being used. Two applications of SSC resins, sealants, and oriented shrink films where EVAs have been traditionally used will be discussed in detail later. Many detailed studies of the thermal properties of SSC technology, homogeneous ethylene–α-olefin copolymers, have been performed by many different institutes [11,16,18]. The first publication showing the effect of increasing comonomer content on the melting points of homogeneous copolymers was U.S. Patent 3,645,992 by Elston of DuPont, Canada [19]. The DSC melting curve of a polyolefin elastomer (POE) at 0.87 g/cm3 density and one melt index is shown in Fig. 18. The glass transition temperature of the POE resin was about ⫺54°C. The POE exhibited a very broad single melting peak with a shoulder and a peak melting temperature of 59°C. The melting begins at temperatures as low as about ⫺40°C. A substantial amount of crystallinity is molten by room temperature. The broad melting range of the POE can be attributed primarily to the intramolecular comonomer distribution. For example, the longer ethylene runs may form a larger lamella crystal with a higher melting point and the shorter one may form a small lamella crystal or even a fringe micelle crystal with a lower melting temperature.

246


FIG. 18

DSC melting curve of ethylene–octene polyolefin elastomer at 0.87 g/cm3, made using CGC technology.

Dynamic Mechanical Properties Dynamic mechanical properties of the homogeneous copolymers as a function of density are illustrated in Figs. 19 and 20. The storage modulus decreases with increasing temperature due to a decrease in crystallinity. The storage modulus decreases sharply near the melting point. The γ-transition peak of all four copolymers and the homopolymer appear at about ⫺120°C in the plot of tan δ versus temperature. In some literature the γ peak was designated as the true glass transition temperature (Tg) of polyethylene. The β-transition peak for the polyolefin elastomer (0.870 g/cm3) appears at about ⫺40°C; for the polyolefin plastomer (0.898 g/cm3), it is about ⫺25°C, and for the 0.920-g/cm3 copolymer, it is about ⫺20°C. The 0.954-g/cm3 polymer has a broad β peak, also at around ⫺20°C. In most cases, the β-transition temperature can be related to the low-temperature performance of ethylene-based polymers and copolymers. For example, below the β-transition temperature, the polymer usually becomes brittle. Therefore, most polyethylene users usually report the β-transition temperature as the practical glass transition temperature. Also, the glass transition temperature measured using DSC matches well with that of the β-transition temperature. For example, as shown in Fig. 18, the DSC Tg of the 0.87-g/cm3 elastomer is about ⫺54°C. The β-transition temperature of the 0.87-g/cm3 elastomer is about ⫺40°C. The β-transition temperature would be higher than the DSC Tg due to the high frequency (10 rad/s) used in the measurement. For the 0.87-g/cm3 elastomer, the α-transition is almost merged with the β-transition. For higher-density SSC polyethylene, distinct α-transitions can also observed. The mechanical properties and deformation behavior of these homogeneous polymers will be discussed further in the next section.


247

FIG. 19 Storage modulus of homogeneous ethylene–octene copolymers and a homopolymer as a function of temperature at 10 rad/s.

FIG. 20

Tan δ of homogeneous ethylene–octene copolymers and a homopolymer as a function of temperature at 10 rad/s.

248


Deformation Behavior and Mechanical Properties of Homogeneous Ethylene–␣-Olefin Copolymers Made with SSC Technology Tensile Properties The stress–strain curves of homogeneous ethylene–octene copolymers and a homopolymer, made using CGC technology, are shown in Fig. 21. The polymers were heated well above the melting point and cooled at 1°C/min. The yield region of copolymers is also enlarged in Fig. 21. The density (crystallinity) of polymer profoundly affected the response to deformation, as evident by the broad spectrum of tensile properties. At high densities, the deformation had characteristics common to many semicrystalline thermoplastics with localized yielding and cold drawing [16]. For copolymers having medium densities (0.902, 0.91, and 0.918 g/cm3), significant strain hardening was observed. For copolymers having low densities (less than 0.885 g/cm3), the moduli were low and the deformation was essentially uniform (homogeneous) and elastomeric. Differences in yield behavior of the four homogeneous polyethylenes at different densities are shown in the photographs in Fig. 22. These photographs were taken at 150% engineering strain. These polymers exhibited deformation behavior ranging from necking at room temperature (density greater than about 0.910 g/ cm3) to uniform deformation (density less than about 0.910 g/cm3). The effect of increasing comonomer content on the tensile deformation behavior and the correlation between the structural classification and the large-scale deformation behavior have been studied and discussed in detail in many publications [16,20–24] and, therefore, will not be further discussed in this article.

FIG. 21 Engineering stress–strain properties of homogeneous ethylene–octene copolymers and a homopolymer, measured at a strain rate of 0.04 s⫺1. (From Ref. 16.)


FIG. 22

249

Photographs of profiles of deformation behavior of Type I to IV homogeneous ethylene– octene copolymers, taken at 150% engineering strain. (From Ref. 16.)

Modulus and yield strength are important properties for ethylene homopolymers and copolymers for many commercial applications, such as packaging film, injection molded containers, and so forth. The slope of the stress–strain curve at very low strain (less than or equal to 2% strain) is measured as the modulus of the polymer. Secant modulus at 1% or 2% strain and Young’s modulus are the ones most commonly used. Moduli of the polyethylenes in the solid state are strongly controlled by polymer crystallinity, crystallite size distribution, and fabrication conditions, which results in orientation. Even with this complexity, there are some ‘‘rules of thumb’’ which can be applied to the modulus. For example, it is generally true that the initial modulus increases with density (crystallinity) of the polymer. There is also a direct relationship between initial modulus and crystallite size. It has been shown [24] that two polyethylenes of the same degree of crystallinity can have very different moduli (up to 100% difference), which is related to a difference in crystallite size. The moduli of a series of homogeneous ethylene–octene copolymers made by CGC technology, several homogeneous ethylene–butene copolymers made by MTC technology, and several conventional LLDPEs made by a Z-N catalyst over a range of polymer densities and comonomer content were measured for comparison purposes [25]. These samples were prepared by compression molding at similar heating and cooling conditions to impart the same heat history on each of the samples. The data are summarized in Fig. 23. It seems that if the polymer samples are prepared under similarly controlled conditions, the Young’s modulus of these polymers is simply a function of the polymer density. However, it should be noted that it is very difficult to produce any ethylene copolymer having a density below about 0.885 g/cm3 using Z-N catalysts. Therefore, data for Z-N heterogeneous

250


FIG. 23 Young’s modulus versus density of ethylene copolymers determined on compression-molded plaques.

LLDPE at below 0.885 g/cm3 are not available for comparison purposes. Based on common knowledge of semicrystalline polymers, the very low-density heterogeneous polymer made by Z-N catalysts, if one can make it, may have a higher modulus than the homogeneous copolymer at similar densities. This is because the heterogeneous polymer, no matter how low the density is, will always have some polymer chains that have a very low level of comonomer incorporated. These polymer chains will crystallize to form larger-sized crystals, which may result in higher modulus than the homogeneous polymer at equivalent density.

Elastic Properties of Polyolefin Elastomers Polyolefin elastomers (POEs), very low-density (crystallinity) homogeneous ethylene–α-olefin copolymers, have a fringed micelle crystal morphology and exhibit a very different deformation behavior as compared to the higher-density copolymers having lamella crystal structures. POEs are now commercially available from DuPont Dow Elastomers, LLC (ENGAGE). The small fringed micelle crystals dispersed in the soft, amorphous matrix act as tie-points to anchor the amorphous chains during deformation and, therefore, result in an elastic recovery upon release of stress. The load/unload property under a tensile deformation mode for a representative octene-based POE (ENGAGE EG-8100, 0.870 g/cm3, 1 MI) is illustrated in Fig. 24. The elastic recovery (permanent set) of various polyethylenes is illustrated in Fig. 25. The slow-cooled (15°C/min) 30-mil-thick compression-molded samples were pulled to desired elongation at 10 in./min (gauge length ⫽ 2 in.). The sample was held at the desired elongation for 30 s. The crosshead was brought back to the initial grip separation at 10 in./min and held for 1 min. The sample was pulled


251

FIG. 24 Load/unload behavior of a representative octene-based polyolefin elastomer (ENGAGE EG8100, 0.870 g/cm3, 1 MI). Strain rate, 2.25 min⫺1; specimens were compression molded and cooled at 15°C/min; test was run at room temperature.

again at 10 in./min until the load rises above zero. Percent permanent set is the percent elongation at which the load rises above zero. The permanent set of homogeneous polyethylenes increased with increasing density. At densities below 0.885 g/cm3, the percent set is low (less than about 20%) at elongations less than 100%. However, the unrecovered strain became much higher beyond the 100% strain.

FIG. 25 Permanent set of various polyethylenes as a function of elongation.

252

Polyolefins Produced by Single-Site Catalyst This phenomenon can hypothetically be explained by a slip-link theory in which the fringed micelle crystals are treated as slippage links. Beyond 100% strain, these crystals started to slip, which resulted in permanent deformation. A detailed discussion of the slip-link model and the use of this model to predict the overall elastic properties of POE has recently been published [22]. New applications for this class of elastic materials were developed commercially for wire and cable insulation, shoe soles, elastic fiber and films, foams, and so forth.

Tie-Molecules in Ethylene–␣-Olefin Copolymers Made by SSC Technology Consideration of the crystalline domains has long dominated research to explain the properties of semicrystalline polymers such as linear low-density polyethylene (LLDPE). Although crystalline domains control low-strain properties of semicrystalline polymers such as modulus and yield stress, it has clearly been established by many researchers that large-strain properties such as stretchability, impact, tear, failure processes, and so forth are also controlled by the amorphous region, particularly by tie-molecules, the amorphous chains that bridge adjacent lamellae [26– 33]. However, development of appropriate structure–property relationships in semicrystalline polymers has been perceived to be hindered by the inability to analytically measure relative tie-chain concentration. In the past, relative tie-chain concentration has been semiquantitatively characterized using techniques such as measurement of the brittle fracture strength [27], infrared dichroism after deformation, and chlorination of films [26]. The relative tie-molecule concentration has also been estimated from chain dimensions and the semicrystalline morphology (topology) of the polymers [28–33]. Tie-molecules in the semicrystalline polymer are critical for enhancing mechanical properties such as environmental stress crack resistance (ESCR), impact, tear, and tensile strength. A schematic diagram of the tie-molecule structure in an ethylene–α-olefin semicrystalline copolymer with a lamellar morphology is illustrated in Fig. 26. A main cause for rejection of a polyethylene chain from a crystal is the presence of imperfections along the chain backbone, which are usually branch points formed by the α-olefin comonomer. Without these branch points (e.g., HDPE), the major part of the polymer chain can possibly be incorporated into the same lamella crystal and, thus, few tie-molecules can be formed. This will result in a polymer with very low mechanical strength. Probabilities for forming tie-molecules in a homogeneous ethylene–α-olefin copolymer made by SSC technology versus a heterogeneous copolymer were estimated [34]. Unlike the heterogeneous ethylene copolymer made by conventional Z-N catalyst technology, the SSC technology polymer has a homogeneous distribution of the α-olefin among the polymer chains. This allows all the polymer chains to crystallize similarly, which results in a narrow crystallite size distribution. The effect of density (crystallinity) on relative tie-chain formation probability and concentration is modeled [35] and the results are illustrated in Fig. 27. The probability of tie-chain formation alone does not reflect the actual tie-chain concentration in semicrystalline poly-


253

FIG. 26 Schematic diagram of tie-molecule structure in an ethylene–α-olefin copolymer.

mers. The relative tie-chain concentration would also depend on the concentration of ‘‘junction points,’’ which would approximately depend on the volume fraction of the crystallinity. The relative tie-chain concentration was obtained from the product of tie-chain probability and the volume fraction of the crystallinity. Figure 27 illustrates the optimum density range and the molecular weight effect on tie-molecule formation. The Huang–Brown model used to calculate probability of tie-chain formation does not take into account the effect of the type of α-olefin comonomer. Intrinsic Elmendorf tear strengths on compression-molded 10-milthick films of various ethylene–α-olefin copolymers were measured [36] and the data are illustrated in Fig. 28. Figure 28 shows that higher α-olefin copolymers (octene and hexene) have much better intrinsic tear strength than butene and pen-

FIG. 27 Relative tie-molecule concentration in a SSC-technology ethylene–octene copolymer at various molecular weights.

254


FIG. 28 Intrinsic tear strength of various SSC-technology ethylene–α-olefin copolymers.

tene copolymers, with octene being the highest of all. The optimum tear strength for all polymers studied is found to be in the density range between 0.89 and 0.92 g/cm3. These experimental results are in good agreement with the tie-chain model.

Applications In previous sections, three key distinctions have been noted between ethylene– α-olefin copolymers produced with single-site catalysts and ethylene–α-olefin copolymers produced with Ziegler–Natta catalysts: (1) an ability to incorporate higher levels of the α-olefin to achieve low polymer density or crystallinity, (2) a uniform comonomer distribution giving a lower melting point at a given density, and (3) a narrow MWD with a Mw /Mn of about 2. The ability to incorporate higher levels of comonomer has made possible new product families (e.g., POP and POE). For example, POEs with densities less than 0.885 g/cm3 are available from SSC, whereas traditional Z-N ethylene–α-olefin copolymers typically are not available below about 0.89 g/cm3 density. The narrow molecular-weight and comonomer distributions contribute to several unique properties, including controllable melting points, reduced extractables, reduced blocking, excellent optics, and excellent mechanical properties. The unique characteristics of the homogeneous ethylene–α-olefin copolymers described make them useful for packaging applications. For example, the ability to control melting and crystallization behavior, the excellent optics, and excellent mechanical strength make POPs an ideal candidate to compete with the traditional high-performance sealants such as EVA copolymers containing 9–18% vinyl acetate and ionomers. Indeed, one important commercial application for plastomers is use as a high-performance sealant in multilayer coextruded or laminated films. Table 1 shows several examples of key properties of plastomers produced with SSC and the resulting packaging benefits.


255

TABLE 1 Benefits of Plastomers in Flexible Packaging Polymer properties Toughness of a linear polyethylene Low sealing temperatures Low extractables Excellent optics High oxygen transmission Elasticity Hydrocarbon composition Thermal stability Moisture insensitivity

Packaging benefits Protection of goods, gauge reduction Faster packaging line speeds Better taste and odor characteristics Package appeal Breathable films High recovery and flexibility Polyolefin compatibility, recycle friendly Processing flexibility Bulk handling, good moisture barrier

Food-Packaging Film Applications The primary distinguishing feature of food packaging is the need to prevent contamination of the food and to maintain or extend the shelf-life. The package acts as a barrier to the environment. Barrier properties will vary depending on the specific type of food being packaged. The barrier of a flexible package can be compromised when (1) an hermetic seal of the package is not obtained during packaging, (2) the seal of the package is not maintained during the packaging, distribution, and sales of the product, or (3) the package itself fails due to inadequate puncture, tear, impact, or abrasion resistance. Seal performance, shrinkage, and abuse resistance are some of the critical performance requirements for many, if not all, of the food packaging applications to be described.

High-Performance Sealants Polyethylenes are widely used as sealants in various packaging applications. Many high-performance films are multilayer structures, where the individual layers perform a specific function(s). High-performance sealants often must meet a number of performance requirements, placing constraints on the polymer design. Beyond parameters like heat-seal-initiation temperature (HSIT), caulkability, seal through contamination, and hot-tack range, consideration is given to parameters like processability during extrusion, machinability (stiffness, coefficient of friction, permeability, appearance, cost, and/or FDA compliance. Types of polyethylenes used as high-performance sealants include EVA, ULDPE, ethylene acrylic acid (EAA) copolymers, ionomers, and, recently, polyolefin plastomers. One of the key attributes of a high-performance sealant is the low melting point, which facilitates the key performance requirement of a low seal-initiation temperature. When plotted against resin crystallinity (Fig. 17), the melting points of EVA resins lie very close to the curve for plastomers produced via SSC technology. Hence, plastomers can potentially be used in various films applications where low-melting-point polymers such as EVAs and ionomers are currently in use (e.g., sealants, oriented shrink films, etc.). Other key attributes of a high-performance sealant are the hot-tack strength, hot-tack window, and heat-seal strength.

256

Polyolefins Produced by Single-Site Catalyst Hot-tack strength is a measure of the force required to separate a semimolten, just-made seal. This test is designed to predict the performance on vertical form, fill, and seal (VFFS) equipment, in which the strength of the seal after it has just been made may be the rate-limiting step in this automated process. High hot-tack strength reduces the likelihood that the food product will break through the seal when being dropped into the forming package. The ultimate hot-tack strength is the highest hot-tack strength achieved over a full range of sealing temperatures. Hot-tack strengths of ethylene–octene plastomers and EVA resins are compared in Fig. 29. EVA polymers exhibit much lower hot-tack strength compared to plastomers. This is due to the very high degree of long-chain branching present in EVA copolymers, which significantly lowers diffusion rates across the molten interface. In the case of EVA resins, an increase in melt index actually increases the hottack strength due to higher diffusion rates [37]. Heat-seal strengths of ethylene–octene plastomers and EVA resins are compared in Fig. 30. Interestingly, at the same melting point, EVA resins exhibit somewhat higher heat-seal initiation temperatures compared to plastomer resins (at melt index of approximately 1 or less). For example, although the melting point of 0.896-g/cm3 plastomer is about 94°C and the melting point of ELVAX 3165 (18% VA EVA) is about 86°C, the 0.896-g/cm3 plastomer exhibited lower seal-initiation temperature. This may due to the high degree of LCB significantly lowering diffusion rates in EVA resins. The ethylene–octene plastomer resins exhibit slightly higher seal plateau strengths compared to EVA resins. Polyolefins plastomers with their low, sharp melting points are thus well suited for high-performance sealants because the ultimate seal and hot-tack strengths of the polymer are achieved at low temperatures [38–41]. One of the applications of high-performance sealants is the box inner liners

FIG. 29

Hot-tack strength of ethylene–octene plastomers made using CGC technology and EVA resins as a function of seal bar temperature. Nylon–EAA–sealant (1/1/1.5 mil)-blown coextrusion film was used.


257

FIG. 30 Heat-seal strength of ethylene–octene plastomers made using CGC technology and EVA resins as a function of seal bar temperature. Nylon–EAA–sealant (1/1/1.5 mil)-blown coextrusion film was used.

for dry foods such as cake mix, cereal, and crackers. Whereas a high-oxygen barrier is critical to preserve the freshness of meat and cheese, a high-moisture barrier is essential for preserving the freshness of dry products. A typical plastic film structure for cereal and cake-mix inner liners is a coextrusion of HDPE with a sealant. The sealants typically used for these applications are ionomers or EVAs containing high levels of vinyl acetate (VA) (i.e., ⱖ18% VA). Again, because a strong, hermetic seal is critical for maintaining the barrier of the package, POPs are another option for replacing ionomers or EVA copolymers (18% VA). A comparison of the other critical sealant performance requirements for a polyolefin plastomer (POP 2), an ionomer (Zn ionomer), and an EVA copolymer (18% VA) are discussed below and summarized in Table 2. POP 2 and EVA copolymers (18% VA) have comparable seal-initiation temperatures (69°C and 72°C, respectively). They are followed by the Zn ionomer, with a seal-initiation temperature of 84°C. Table 2 shows that POP 2 provides the highest ultimate hot-tack strength of the three products, followed by the Zn ionomer and the EVA copolymer. Note that the 18% VA copolymer provides very low ultimate hot-tack strength due to very high levels of long-chain branching retarding diffusion across the interface and is generally not suitable for the packaging of heavy products, such as cake mixes.

Oriented Shrink Films Oriented shrink films can be differentiated from conventional hot-blown shrink films by their fabrication processes. The conventional blown shrink-film method involves extrusion of the molten polymer through an annular die at a certain drawdown ratio and a certain blowup ratio and then cooling the bubble via ambient or chilled air. Orientation of the film takes place in a completely molten state and is

258

Polyolefins Produced by Single-Site Catalyst TABLE 2 High-Performance Sealants for Box Inner Liners Resin

POP

Melt index (dg/min) Density (g/cm3) 2.0-Mil monolayer blown-film properties Puncture (ft-lb./in.3) Tear Resistance (g) Haze (%) Odor intensity rating Water vapor transmission rate (g mil/100 in.2 day atm) 1.5-Mil HDPE/0.5-mil sealant, blown coextrusion Heat-seal initiation (°C) Ultimate hot-tack strength (N/in.)

1.6 0.8965 270 300 1.3 1.2

EVA (18% VA)

Zn Ionomer 1

0.8 Not available

5.5 Not available

222 118 1.5 Not available 6.2

104 190 2.9 2.7 Not available

72 3.4

84 4.4

2.0

69 7.5

primarily in the machine direction, due to the drawdown ratio, and secondarily in the cross direction, as a result of the blowup ratio. In contrast, the oriented shrinkfilm process can be described generically by the following steps: Extrusion–quenching–reheating–baxially stretching–cooling The reheating step involves temperatures just above the glass transition temperature in the case of amorphous polymers, or below the peak melting temperature in the case of semicrystalline polymers. There are two basic methods to produce biaxially oriented films. The first method is the tenter frame process and the second method is often referred to as the ‘‘doublebubble’’ or ‘‘trapped-bubble’’ process. More in-depth descriptions of these processes can be found elsewhere [42,43]. Orientation by these processes takes place in a semimolten state (i.e., at a temperature below the melting point of the semicrystalline polyolefin). As a result, a significantly higher degree of orientation is obtained compared to hot-blown shrink films. These orientation processes produce high-tensile, high-modulus films with high shrink and shrink tension, as well as excellent clarity [42,43]. Biaxially oriented films made via the ‘‘doublebubble’’ or ‘‘trapped-bubble’’ process are used to produce packages for very large, subprimal cuts of meat after slaughter. The packages are often referred to as barrier bags. The subprimal cuts of meat are inserted into a barrier bag that is generally sealed at one end, the air evacuated, and then the bag heat sealed. The package is then heated, usually via hot water, to obtain shrinkage of the film around the meat [44]. Typical barrier bag structures range from three to five layers and minimally contain an abuse/shrink layer, an oxygen barrier layer such as a poly(vinylidene chloride) (PVDC) or ethylene vinyl alcohol copolymer (EVOH), and a sealant layer. In addition, adhesive layers may be used to tie the structure together. The use of LLDPEs and ULDPEs alone or as blends with EVA copolymers in both the abuse and sealant layers have been taught in the patent literature [44,45]. The broad comonomer distributions and melting ranges of LLDPEs and ULDPEs create a sufficiently broad orientation window for the double-bubble process. Fur-


FIG. 31

259

Hot-water shrinkage at 90°C of films oriented on a T.M. Long stretcher under isothermal conditions at a 4.5 ⫻ 4.5 draw ratio at 5 in./s.

thermore, the excellent toughness of these linear polymers provides the high level of abuse resistance needed to prevent punctures from the bones in subprimal meats and to sustain impact abuse during the distribution process. ULDPE resins offer improvements in shrinkage over higher-density LLDPE resins, due to their reduced crystallinity. Further improvements in shrinkage can be obtained using POPs. POPs can offer low-temperature shrinkage improvements over ULDPEs at the same density because of their lower melting points. Figure 31 illustrates the improved shrinkage at 90°C for POPs at a density below about 0.908 g/cm3 [46]. In addition, it is assumed that the improved sealability and abuse-resistance properties for POPs in conventional blown films also translate to improvements in oriented shrink films. As mentioned earlier, blends of EVA with ULDPE or LLDPE have been taught for use in barrier bag structures. A particular disadvantage of these blends is the poor optics and reduced abuse resistance. As shown in Table 3, the POP/LLDPE blend provides significantly better optics and potentially better toughness in an oriented shrink film than the EVA/LLDPE blend. Both the POP and EVA have compaTABLE 3

Optics of Oriented Films

Blend ratio Haze (%) Shrinkage at 90°C (%)

EVA (12% VA)/ LLDPE

POP*/LLDPE †

30%/70% 7.6 16

30%/70% 0.5 16.5

Note: Films oriented on T.M. Long stretcher under isothermal conditions at a 4.5 ⫻ 4.5 draw rate at 5 in./s. * POP is an ethylene–octene copolymer having a melt index of 0.9 and density of 0.898 g/cm3. † LLDPE is an ethylene–octene copolymer having a melt index of 1.0 dg/ min and density of 0.920 g/cm3.

260

Polyolefins Produced by Single-Site Catalyst rable melting points and crystallinity. The poor optics of the EVA/LLDPE blend are believed to be due to the partial immiscibility of these two polymers [47]. Other important applications of homogeneous polyethylenes made using SSC technology include fresh-cut produce packaging [48–50], high-performance films [51–54], batch inclusion bags [55], cling layer in stretch films [56,57], extrusion coating [58,59], and impact modification of polypropylenes [60–65].

Conclusion and Summary Development and implementation of SSC technology have greatly expanded the product range of polyolefins. These homogeneous polymers have narrow molecular-weight and composition distributions, resulting in enhanced mechanical, optical, and heat-seal properties. Because of the simple and predictable molecular structure of this class of polyolefin copolymers, building a model to predict properties for these polymers became feasible. This allows the materials scientists to design products by a molecular-architecture approach, using property and performance requirements as guidelines. With the advancement of the SSC technology for the production of polyolefin copolymers with well-defined molecular structures, the authors expect the polyolefin industry can take great advantage of this technology to design unique polymers and fulfill their customer needs in the future. Homogeneous polyolefins provide excellent utility in numerous packaging applications, ranging from retail meat packages to industrial stretch films. These resins may be fabricated into packaging films or containers via blown- and cast-film methods, biaxial oriented film processes, extrusion coating, lamination, and injection molding. In packaging applications, the resins based on CGC technology have successfully replaced and/or upgraded EVA, ionomer, LDPE, LLDPE, and ULDPE.

Acknowledgments The authors of this article would like to express their sincere appreciation to Professor Eric Baer, Professor Anne Hiltner of Case Western Reserve University, and Pradeep Jain, Kalyan Sehanobish of The Dow Chemical Company. They provided many useful data in this article.

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