Encyclopedia of Chemical Processing and Design: Project Progress Management to Pumps 082472495X, 9780824724955

Citation preview

Encyclopedia of Chemical Processing and Design 45

EXECUTIVE EDITOR JOHN J. McKETTA

The University of Texas at Austin Austin, Texas EDITORIAL ADVISORY BOARD

LYLE F. ALBRIGHT

JAMES R. FAIR

JOHN HAPPEL

Purdue University Lafayette, Indiana

Professor of Chemical Engineering The University of Texas Austin, Texas

Columbia University New York, New York

ERNEST E. LUDWIG

R. A. McKETTA

Ludwig Consulting Engineers, Inc. Baton Rouge, Louisiana

Chemical Engineer Purvin and Gertz, Inc. Houston, Texas

Encyclopedia of Chemical Processing and Design_______ John J. McKetta

e x e c u t iv e e d it o r

Project Progress Management to Pumps

MARCEL DEKKER, INC.

NEW YORK • BASEL • HONG KONG

Library of Congress Cataloging in Publication Data Main entry under title: Encyclopedia of chemical processing and design. Includes bibliographical references. 1. Chemical engineering— Dictionaries Technical— Dictionaries. I. M cKetta, John J. II. Cunningham , W illiam Aaron. Tp9.E66 660.2 '8 '0 0 3 ISBN: 0-8247-2495-X

2.

Chemistry,

75-40646

Pages 246-535 were originally published as Pumps for Chemical Processing by J. T. McGuire, ° 1990, Marcel Dekker, Inc., New York. COPYRIGH T © 1993 by M ARCEL DEKKER, INC. ALL RIGHTS RESERVED. N either this book nor any part may be reproduced or transmitted in any form or by any m eans, electronic or mechanical, including photocopying, m icrofilming, and re­ cording, or by any information storage and retrieval system , without permission in w riting from the publisher. M ARCEL DEKKER, INC. 270 M adison Avenue, New York, New York, 10016 Current printing (last digit): 10 9 8 7 6 5 4 3 2 1 PRIN TED IN TH E UN ITED STATES OF AM ERICA

International Advisory Board RAY C. ADAM

LUCIANO BENINCAMPI

NICHOLAS P. CHOPEY

Former Chairman of the Board N. L. Industries, Inc. New York, New York

Manager of Public Relations CTIP— Compagnia Tecnica Industrie R om e, Italy

Editor-in-Chief Chemical Engineering Magazine McGraw-Hill, Inc. New York, New York

CARL W. ALBERS

LLOYD BERG

Senior Process Engineer M. W. Kellogg Houston, Texas

Professor Department of Chemical Engineering Montana State University Bozeman, Montana

M. A. ALLAWALA Managing Director National Refinery Ltd. Karachi, Pakistan

HAMED H. AMER Chairman Agiba Petroleum Co. Cairo, Egypt

R. G. ANTHONY Professor, Department of Chemical Engineering Texas A & M University College Station, Texas

H. J. AROYAN Former Vice President Chevron Research Company Richmond, California

F. SID ASKARI President Technolog, Inc. Engineering and Industrial Consultants Tehran, Iran

DONALD L. BAEDER Former Executive Vice President— Science and Technology Occidental Petroleum Corporation Los Angeles, California

Wm. A. BAILEY, Jr. Former Director, MTM Process Research and D evelopm ent Lab Shell D evelopm ent Company H ouston, Texas

TRAVIS W. BAIN Vice President National Sales, Inc. Jackson, Mississippi

GAREN BALEKJIAN C. F. Braun Arcadia, California

CESAR BAPTISTE Vice President Petroleos Mexicanos M exico City, Mexico

NEIL S. BERMAN Professor of Chemical Engineering Engineering Center Arizona State University Tempe, Arizona

D. J. BLICKWEDE Former Vice President and Director of Research Bethlehem Steel Corp. Bethlehem, Pennsylvania

M. J. P. BOGART Fluor Engineers and Constructors, Inc. Santa Ana, California

Z. D. BONNER Vice Chairman of the Board Tesoro Petroleum Corp. San Antonio, Texas

JOSEPH F. BOSICH Bosich Consultants Humble, Texas

WILLIAM H. BOSLER President Texas Consultants, Inc. Houston, Texas

ARCHIE BROODO President A ID , Inc. Dallas, Texas

ARTHUR W. BUSCH Environmental Engineer Consultant Dallas, Texas

ROBERT C. BUTLER Administrative Assistant and Planning Manager, Petroleum Chemicals Division E. I. du Pont de Nemours and Co. Wilmington, Delaware

J. MORSE CAVENDER President The Mactan Company Dusseldorf, Federal Republic of Germany

LEON R. MARTINEZ BASS

PRAMOTE CHAIYAVICH

Sales Manager— Northern M exico Zincamex, S. A. Saltillo, Mexico

Chief Technologist The Tahi Oil Refinery C o., Ltd. Bangkok, Thailand

ROBERT O. BATH IANY

S. D. CHELLAPPAN

Technical Planner Weyerhauser Company Tacoma, Washington

Process Engineering Manager Occidental Chemical Corporation Houston, Texas

FRANK CHRENCIK Vulcan Materials Co. Birmingham, Alabama

R. JAMES COMEAUX Vice President American Petrofina Dallas, Texas

C. W. COOK Chairman, Executive Committee General Foods Corp. White Plains, New York

CHARLES F. COOK Vice President Research and D evelopment Phillips Petroleum Bartlesville, Oklahoma

EARL J. COUCH Research Associate Mobil Research and Development Corp. Dallas, Texas

JAMES R. COUPER Professor Department of Chemical Engineering University of Arkansas Fayetteville, Arkansas

HORACE R. CRAWFORD Senior Staff Scientist CONOCO Corp. Houston, Texas

ORAN L. CULBERSON Chemical Engineer Oak Ridge National Lab Chemical Technology Division Oak Ridge, Tennessee

DONALD A. DAHLSTROM Vice President, Research and D evelopm ent Process Equipment Group Envirotech Corp. Salt Lake City, Utah

PERRY P. DAWSON Production Engineer D ow Chemical Co. Freeport, Texas

ELBERT M. DeFOREST Former Director of Technology, Chemicals and Metals Vulcan Materials Co. Wichita, Kansas

ROBERT G. DENKEWALTER Corporate Vice President Technology Allied Corp. Morristown, New Jersey

iii

International Advisory Board

iv

J. P. de SOUSA

RALPH T. FERRELL

OM P. GOYAL

Publisher Chemical A ge of India Technical Press Publication Bombay, India

Senior Vice President, Corporate D evelopment Vista Chemical Company Houston, Texas

Technagement Consultant New Bombay, India

JAMES D. D'lANNI

LOUIS FEUVRAIS

Former Director of Research The Goodyear Tire and Rubber Co. A kron, Ohio

Directeur General Ecole Nationale Superieure D ’Arts et Metiers Paris, France

JUAN M. DIAZ Production General Manager Rohm and Haas M exico, S. A . C. V. M exico City, Mexico

WERNER DIMMLING Dipl-Chemist Friedrich Uhde GmbH Dortmund, Federal Republic of Germany

S. W. DREW Executive Director MCMC Technical Operations Merck & C o., Inc. Rahway, New Jersey

BARRETT S. DUFF Barrett S. D uff and A ssociates South Pasadena, California

P. K. DUTTA Project Manager Chemical and Metallurgical Design Company, Private Ltd. N ew D elhi, India

WILLIAM F. EARLY Vice President Stone & Webster Environmental Services H ouston, Texas

WALTER EMRICH Consultant Teterboro, New Jersey

E. FREDERICO ENGEL Member o f the Board of Management Chemische Werke Hiils A G Marl, Federal Republic o f Germany

P. E. G. M. EVERS Operations Manager A nzo Salt Chemical Delfzyl D elfzyl, The Netherlands

ALEXANDRE EVSTAFIEV Director, Division o f Technological Research and Higher Education U N ESC O — Paris Paris, France

GERALD L. FARRAR President Farrer A ssociates Tulsa, Oklahoma

F. M. FARRELL Technical Director 3M Company St. Paul, M innesota

C. SHULTS FAULKNER President C. S. Faulkner, Inc. H ouston, Texas

R. A. FINDLAY Former Director, Fuels and Lubricants, Research and D evelopm ent Phillips Petroleum Company Bartlesville, Oklahoma

DALE FRIDLEY Manager, Intermediates Technology Division Exxon Chemical America Baton Rouge, Louisiana

ROBERT H. FRITZ President Loss Control Consultants, Inc. Alvin, Texas

GARY L. FUNK Director, Advanced Process Control Technology Brown & R oot/IT I Division Houston, Texas

BILL F. GALLOWAY Plant Manager Quantum USI Division Port Arthur, Texas

DONALD E. GARRETT President Saline Processors Ojai, California

L. W. GARRETT, Jr. President Garrett Associates, Inc. San M ateo, California

ROY D. GERARD General Manager Westhollow Research Center Shell Developm ent Company Houston, Texas

ION GHEJAN Department of Chemical Engineering Institute of Petroleum, Gas, and Geology Bucharest, Romania

JIM GILLINGHAM General Manager, Process Engineering Diamond Shamrock San Antonio, Texas

WILHELM GRAULICH Director, Manager, Rubber Division Bayer A G Leverkusen, Federal Republic of Germany

E. HENRY GROPPE Groppe, Long, & Littell H ouston, Texas

GIANFRANCO GUERRERI INGECO Altech Group Societa per A zioni Con Sede in Milano Milan, Italy

KENNETH M. GUTHRIE Cost Consultant Marina D el R ey, California

NORMAN HACKERMAN Former President Rice University H ouston, Texas

VLADIMIR HAENSEL Vice President, Science and Technology Universal Oil Products Co. D es Plaines, Illinois

HENRY E. HALEY Vice President Arthur D . Little, Inc. Cambridge, Massachusetts

R. L. HARVEL Project Manager D ow Chemical International Ltd. T okyo,Japan

J. W. HAUN Former Vice President and Director of Engineering General Mills, Inc. M inneapolis, Minnesota

TERUAKI HIGUCHI President Japan Fody Corp. Osaka, Japan

JOHN R. HILL, Jr. President and Chief Executive Officer Gifford-Hill & C o., Inc. Dallas, Texas

PAUL E. HIME Former Vice President Operation & Technical Hoechst Celanese Chemical Group D allas, Texas

HAROLD L. HOFFMAN

B. GENE GOAR

Editor Hydrocarbon Processing H ouston, Texas

Goard, Allison, and Associates, Inc. Tyler, Texas

NORBERT IBL

MARCEL GOLDENBERG SAM IN Corp., Inc. New York, New York

Professor Eidg. Techn. Hochscule Zurich Techn.— Chemie Zurich, Switzerland

International Advisory Board

v

RUBEN F. INGA

W. S. LANIER

BRYCE I. MacDONALD

President Confederacion Interamerican de Ingeniera Quimica Lima, Peru

Project Manager Seadrift Expansion Projects Union Carbide Corp. Houston, Texas

Manager, Environmental Engineering General Electric Company Fairfield, Connecticut

JAMES R. JOHNSON

CLARK P. LATTIN, Jr.

Former Executive Scientist and Director, Advanced Research Programs Laboratory 3M Company, Central Research Labs Saint Paul, Minnesota

Former President The M. W. Kellogg Company Houston, Texas

Professor School of Chemical Engineering Oklahoma State University Stillwater, Oklahoma

NAJI A. KADIR President Scientific Research Council Baghdad, Iraq

JOHN E. KASCH Former Vice President Standard Oil Indiana Escondido, California

RAPHAEL KATZEN Managing Partner Ralph Katzen Associates Cincinnati, Ohio

JOHN J. KELLY

ISIDORO LAZARRAGA-LEANZA

KLAUS MAI

Chief of Engineering and Control Empresa Nacional del Petroleo Vina del Mar, Chile

Former President Shell D evelopm ent Houston, Texas

JEAN Le BRETON

STANLEY D. MARTS

Managing Director Elf Aquitaine Indonesie Jakarta, Indonesia

Supply Specialist Shell Oil Company Houston, Texas

IRV LEIBSON Vice President Bechtel Corp. San Francisco, California

PIERRE Le PRINCE

Department of Chemical Engineering University College, Dublin Dublin, Ireland

Director of Refining and Engineering Center Instiut Francaise de Petrole Malmaison, France

HENNO KESKKULA

C. E. LETSCHER

Research Fellow Chemical Engineering Department The University of Texas at Austin Austin, Texas

Caltex Petroleum Company New York, New York

C. J. LIDDLE

0. P. KHARBANDA

White Young & Partners Ltd. Herts, England

O. P Kharbanda & Associates Cost and Management Consultants Bombay, India

NORMAN N. LI

WLODZIMIERZ KiSIELOW Professor of Petroleum Technology, Director of Research Department of Petroleum and Coal Centre of Polish Academy of Sciences Krzywoustego, Poland

ROBERT A. KLEIN President and Chief Executive Officer Continental Controls, Inc. Houston, Texas

MOHAN SINGH KOTHARI Chief Consultant Punjab Industrial Consultancy Organisation Chandigarh, India

G. R. KRUGER President Semarck, Inc. Houston, Texas

A. P. KUDCHADKER Professor of Chemical Engineering and D ean of Student Affairs Indian Institute of Technology, Kanpur Kanpur, India

R. N. MADDOX

Director, Chemical & Process Technology Allied Signal Engineered Materials Research Center D es Plaines, Illinois

DAVID C. K. LIN Senior Engineer Owens Corning Fiberglas Corp. Newark, Ohio

CHARLES E. LOEFFLER Technical Manager Celanese Chemical Company Pampa, Texas

T. N. LOLADZE Vice-Rector, Professor of the Georgian Polytechnic Institute Tbilisi, USSR

STANLEY L. LOPATA Chairman of the Board Carboline Company Saint Louis, Missouri

PHILIPS S. LOWELL Chemical Engineer Consultant Austin, Texas

RALPH LANDAU

W. D. LUEDEKE

Former Chairman Halcon International, Inc. New York, New York

Former Planning Manager E. I. du Pont de Nemours Wilmington, Delaware

F. DREW MAYFIELD Drew Mayfield & Associates Baton R ouge, Louisiana

GUY McBRIDE Former President Colorado School of Mines Golden, Colorado

CLYDE McKINELY Former D irector,Allentown Labs Air Products and Chemicals, Inc. A llentown, Pennsylvania

RICARDO MILLARES President Papel Satinado, S. A. M exico City, Mexico

ROBERT L. MITCHELL Former Vice Chairman of the Board Celanese Corp. New York, New York

RICHARD MOLLISON General Manager Colpapel, S. A . Pereira, Columbia

DONALD D. MULRANEY C. F. Braun Co. Alhambra, California

CARLOS EPSTEIN MURGUIA General Manager and President of the Board Industrias Guillermo Murguia, S. A. Naucalpan, M exico

TAKAYUKI NATE Plastics Sales Department Tonen Petrochemical Co. Ltd. T okyo,Japan

JAMES K. NICKERSON Research Associate Esso Research and Engineering Company Summit, New Jersey

ALEX G. OBLAD Distinguished Professor of Chemistry Mining, and Fuels Engineering University of Utah Salt Lake City, Utah

International Advisory Board

vi

H. E. O'CONNEL

JEFFREY P. REILLY

GERT G. SCHOLTEN

Former President Tenneco Chemicals Inc. Houston, Texas

Manager Technology Acquisition & D evelopm ent A BB Lummus Crest Inc. Houston, Texas 79570

Managing Director Edeleanu Gesellschaft mbH Frankfurt/Main. Federal Republic of Germany

AURELIO REITER

Vice President Business Developm ent & Technology Exxon Chemical Company Darien, Connecticut

ERNEST 0. OHSOL Consultant Ohsol Technical A ssociates Crosby, Texas

I. 0. OLADAPO Dean of Engineering University of Lagos Lagos, Nigeria

GORDON F. PALM President Gordon, F. Palm & Associates Lakeland, Florida

F. F. PAPA-BLANCO Advisor o f Educational Technology Instituto Latino Americano de la Communicacion Educativa Mexico City, Mexico

W. M. PARRIS Technical Consultant, TIMET Las Vegas, Nevada

DILIP M. PATEL Manager of Process Design & Technology John Brown E & C, Inc. Houston, Texas

MARCELLO PICCIOTTI Technical Promotion Manager TechniPetrol-Rome Rom e, Italy

THOMAS C. PONDER Petrochemicals Editor Hydrocarbon Processing Houston, Texas

R. G. H. PRINCE Professor, Head o f Department Chemical Engineering University of Sydney Sydney, Australia

HUGH S. PYLANT Project Manager The Pace Consultants, Inc. Houston, Texas

EDWIN L. RAINWATER D ow Chemical U SA Texas Operations Industrial Chemicals Division Freeport, Texas

J. S. RATCLIFFE Professor of Chemical Engineering University of New South Wales Kensington, Australia

Former Research Manager of Esso Standard Italiana Roma-Italy R om e, Italy

LARRY RESEN

M. L. SHARRAH

Larrry Resen Associates W ilton, Connecticut

Former Senior Vice President Continental Oil Company Stamford, Connecticut

H. KEN RIGSBEE

JOHN W. SHEEHAN

Project Manager Phillips 66 Natural Gas Company Houston, Texas

Vice President, Manufacturing and Marketing Champlin Petroleum Company Kerrville, Texas

FRANK S. RIORDAN, Jr. Director, Technology Planning Monsanto Textiles Company Saint Louis, Missouri

DENNIS F. RIPPLE Technical Manager,Process Technology Hoechst Celanese Corporation Corpus Christi, Texas

LOUIS R. ROBERTS Director, Planning and Source Evaluation Texas Air Control Board Austin, Texas

RICCARDO ROBITSCHEK Direttore Divisione Resine Societa Italiana Resine M ilano, Italy

ROBERTO RODRIGUEZ INTEVEP Caracas, Venezuela

GERHARD ROUVE Director of the Institute for Water Resources D evelopm ent Technical University Aachen A achen, Federal Republic of Germany

PIERRE SIBRA Designer Esso Engineering Services Ltd. Surrey, England

PHILIP M. SIGMUND Professor of Chemical Engineering University of Calgary Alberta, Canada

ARTHUR L. SMALLEY, Jr. President Matthew Hall Inc. Houston, Texas

CARL I. SOPCISAK Technical Consultant Synthetic Fuels Wheat Ridge, Colorado

PETER H. SPITZ President Chemicals Systems Inc. New York, New York

SAM STRELZOFF Consultant Marlboro, Vermont

MARK B. STRINGFELLOW President & Chief Executive Officer Environmental Control Group, Inc. Maple Shade, New Jersey

JOHN H. SANDERS

Y. S. SURY

Vice President and Assistant General Manager Eastman Chemicals Division Eastman Kodak Company Kingsport, Tennessee

CIBA-Geigy Chemical Corp. Saint Gabriel, Louisiana

HIDESHI SATO General Manager Technical Information Office Technical D evelopment Department Nippon Steel Corp. Tokyo, Japan

MICHAEL W. SWARTZLANDER Staff Engineer Union Carbide Corp. South Charleston, West Virginia

T. SZENTMARTONY A ssociate Professor Technical University Budapest Budapest, Hungary

M. TAKENOUCHI

FRANCIS E. REESE Former Vice President and Managing Director International Monsanto Company Saint Louis, Missouri

DOUGLAS M. SELMAN

GEORGE E. SCHAAL Manager, Research and D evelopm ent Produits Chimiques Ugine Kuhlmann Pierre-Benite, France

General Manager of Manufacturing Department Maruzen Oil C o., Ltd. T okyo,Japan

vii

International Advisory Board

VLADIMIR TEPLYAKOV Head of Membrane Research Center A. V. Tochiev Institute of Petrochemical Synthesis The USSR Academy of Sciences Moscow, Russia

SOONTHORN THAVIPHOKE Managing Director S. Engineering Services C o., Ltd. Bangkok, Thailand

ROBERT S. TIMMINS Core Laboratory Aurora, Colorado

A. A. TOPRAC President Interchem-Hellas Athens, Greece

YORGI A. TOPRAKCLOGLU

JUAN JOSE URRUELA VILLACORTA

PAUL B. WEISZ

Ingeniero Fabrica de Jabon “La Luz, S. A .” Guatemala

Distinguished Professor Chemical and Bio-Engineering University of Pennsylvania Philadelphia, Pennsylvania

S. P. VOHRA

JACK W. WESTERFIELD

Managing Director Bakelite Hylam, Ltd. Bombay, India

Manager, Project Engineering Diamond Shamrock San A ntonio, Texas

A. L. WADDAMS

D. L. WILEY

Former Manager, Marketing Services Division BP Chemicals International Ltd. London, England

Former Senior Vice President U nion Carbide Corp. Danbury, Connecticut

T. J. WALKER Former Production Manager D ow Chemical Europe S. A . Zurich, Switzerland

JACK C. WILLIAMS Former Vice President T exaco,In c. Houston, Texas

MASAMI YABUNE

Chairman of the Board of Directors Marshall Boya ve Vernik Sanayii A . S. Istanbul, Turkey

J. C. WALTER, Jr.

GEORGE H. UNZELMAN

THEODORE WEAVER

LEWIS C. YEN

President Hy Ox, Inc. Fallbrook, California

Director of Licensing Fluor Corporation Los A ngeles, California

Manager, Technical Data M. W. Kellogg Company Houston, Texas

HERNANCO VASQUEZ-SILVA

ALBERT H. WEHE

STANELY B. ZDONIK

President Hernando Vasqez & A ssociates, Ltd. Bogota, Columbia

Chief, Cost and Energy U. S. Government Raleigh, North Carolina

M. A. VELA

GUY E. WEISMANTEL

Vice President and Manager Process Department Stone and Webster Engineering Corp. Boston, Massachusetts

President VELCO Engineering, Inc. H ouston, Texas

President Weismantel International Houston, Texas

J. C. Walter Interests Houston, Texas

Section Head, Technical Section Tonen Petrochemical C o., Ltd. T okyo,Japan

Contributors to Volume 45 Wallace P. Bolen M ineral C om m odity Specialist, U .S. Bureau o f M ines, W ashington, D .C .: Pum ice and Pum icite Supply-Dem cm d R elationships

Carl L. Elmore Vice P resident, K am yn, Inc., G lens Falls, New York: Pulp and Pulping George Georgiou, Ph.D. A ssociate Professor o f C hem ical E ngineering, The U ni­ versity o f Texas at A ustin, A ustin, Texas: Protein Engineering

Jeffrey A. Hubbell, Ph.D. A ssociate Professor o f C hem ical E ngineering, The U ni­ versity o f Texas at A ustin, A ustin, Texas: Protein Engineering

A.E. Kerridge The M .W .

K ellogg

Com pany,

H ouston,

Texas:

Project

P rogress

M anagem ent

David I. Limb David Lim b A ssociates, S tockport, England: Propane Recovery, RefluxExc'hange P rocess

Pat McCann M anager o f G as Plants, A m erada Hess C o rp o ratio n , H ouston, Texas: Propane Recovery, Ryan/H olm es P rocess

J.T. McGuire D resser Industries, Inc., H untington P ark, C alifornia: Pumps Robert J. McKee Principal Engineer, M echanical and Fluids E ngineering D ivision, S outhw est R esearch System s

Institute, San A ntonio, Texas: Pulsation in G as M etering

Yuv R. Mehra Director, New Product D evelopm ent, El Paso H ydrocarbons Com pany, O dessa, Texas: Propane Refrigeration System s; Propylene Refrigeration System s

John V. O'Brien Vice P resident, Process T echnology G roup, Process System s Inter­ natio n al, Process

Inc.,

W estborough,

M assachusetts:

P ropane

Recovery,

Ryan/H olm es

Donald I. Orenbuch A R C O C hem ical Com pany, N ew tow n Square, Pennsylvania: Propylene Oxide and G lycol

Joseph F. Pilaro Vice P resident, C hem icals, M obil Saudi A rabia, In c., K ingdom o f Saudi A rabia: Projecting C ost via E conom etric M odeling

James M. Ryan, Sc.D. E ngineering Fellow, Koch Industries, In c., W ichita, Kansas: P ropane Recovery, Ryan/H olm es P rocess

Cecil R. Sparks Vice President (retired), A pplied Physics D ivision, S outhw est R esearch Institute, San A ntonio, Texas: Pulsation in Centrifugal C om pressors; Pulsation in C entrifugal Pum ps; Pulsation in G as M etering System s

Terry R. Tomlinson C h ief Process E ngineer, C ostain O il, G as and P rocess, M an­ chester, E ngland: Propane Recovery, Reflux-Exchange P rocess

Francis H. Tung N ational D istillers and C hem ical C o rp o ra tio n , New York, New York: Projecting C ost via E conom etric M odeling

J.R. Valbert A R C O C hem ical Com pany, N ew tow n Square, Pennsylvania: P ropylene Oxide and G lycol

John Veranth A M A X Scheduling

M agnesium

C o rp o ratio n ,

Salt

Lake

City,

Utah:

P roject

Contributors to Volume 45

J.C. Wachel M anager, A pplied M echanics, S outhw est R esearch Institute, San A ntonio, Texas; (currently) P resident, E ngineering D ynam ics, Inc., San A ntonio, Texas: P ul­ sation in Centrifugal Pumps

J.G. Zajacek A R C O C hem ical C om pany, N ew tow n Square, Pennsylvania: P ropylene Oxide and G lycol

Adam Zanker, Ch.E., M.Sc. Senior R esearch Engineer, Oil R efineries, L td ., H aifa, Israel: Pulp Slurries, D ensity o f

Contents of Volume 45 Contributors to Volume 45

ix

Conversion to SI Units

xiii

Bringing Costs up to Date

xv

Project Progress Management A. E. K erridge

1

Project Scheduling John Veranth

17

Projecting Cost via Econometric Modeling Francis H. Tung and Joseph F. Pilaro

34

Propane Recovery, Reflux-Exchange Process T erry R. Tom linson and David I. Lim b

43

Propane Recovery, Ryan/Holmes Process Jam es M. R yan, Pat M cC ann, and John V. O ’B rien

59

Propane Refrigeration Systems Yuv R. M ehra

76

Propylene Oxide and Glycol J. R. V albert, J. G. Z ajacek, and D onald I. O renbuch

88

Propylene Refrigeration Systems Yuv R. M ehra

128

Protein Engineering G eorge G eorgiou and Jeffrey A. Hubbell

142

Pulp and Pulping C arl L. Elm ore

177

Pulp Slurries, Density of Adam Z anker

190

Pulsation in Centrifugal Compressors Cecil R. Sparks

192

Pulsation in Centrifugal Pumps Cecil R. Sparks and J. C. W achel

215

Pulsation in Gas M etering Systems Cecil R. S parks and R obert J. M cK ee

229

Pumice and Pum icite Supply-Demand Relationships W allace P. Bolen

241

Pumps J. T. M cG uire

246

Conversion to SI Units

X

©

oo oo oo

©

X

©

1.60 x 10‘ 19 2.96 x 1 0 "5 0.305 2.01 x 102 4.404 x 1 0 -3 3.785 x 10"3 1.183 x 10"4 6.48 x 1 0 "5 X

1

o

7.457 x 102 9.81 x 103 7.46 x 102 50.80 ©

X

45.36 to L*

kilogram (kg) meter (m) newton/square meter (N /m 2) newton/square meter (N /m 2) newton (N)

x 10‘ 2 10’ 7 x 10"3

©

kilogram (kg)

x 103

3.697 x 1 0 '6

p

hundred weight (short) inch inch mercury inch water kilogram force

joule (J) joule (J) cubic meter (m3) 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)

4.187 4.190 4.184 1.333 98.06 0.457 1.745 1.0 x 1.772

X

gill (U.S.) grain gram horsepower horsepower (boiler) horsepower (electric) hundred weight (long)

newton (N)

0.159 1.055 x 103 1.056 x 103 1.054 x 103 3.52 x 1 0 '2

p

dram (U.S. fluid) dyne electron volt erg fluid ounce (U.S.) foot furlong gallon (U.S. dry) gallon (U.S. liquid)

kilogram (kg) kilogram (kg) cubic meter (m 3)

©

cubit degree (angle) denier (international) dram (avoirdupois) dram (troy)

joule (J) joule (J) newton/square meter (N /m 2) newton/square meter (N /m 2) meter (m) radian (rad) kilogram/meter (kg/m)

X

calorie (mean) calorie (thermochemical) centimeter of mercury centimeter of water

1.013 x 10s p

bushel calorie (International Steam Table)

joule (J) joule (J) joule (J) cubic meter (m 3) joule (J)

o

4.046 x 103 © ©

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)

X

acre angstrom are atmosphere bar barrel (42 gallon) Btu (International Steam T^ble) Btu (mean) Btu (thermochemical)

p

M ultiply by

X

To

p

To convert from

3.386 x 103 2.49 x 102 9.806

xiv

C o n v e rs io n to SI U n its

To convert from

To

M ultiply by

kip knot (international) league (British nautical) league (statute) light year liter micron

newton (N) meter/second (m/s) meter (m) meter (m) meter (m) cubic meter (m 3) meter (m)

4.45 x 103 0.5144 5.559 x 103 4.83 x 103 9.46 x 1015

mil mile (U.S. nautical) mile (U.S. statute) millibar millimeter mercury

meter (m) meter (m) meter (m)

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

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 (m2/s) kilogram (kg) kilogram (kg) kilogram (kg) newton/square meter (N /m 2) meter (m)

0.001

1.0 x 10‘6 2.54 x 10‘6 1.852 x 103 1.609 x 103 100.0

1.333 x 102 79.58 0.278 2.835 x 10 2 3.11 x 10"2 2.96 x 10'5 1.0

8.81 x 10"3 1.555 x 10‘3 5.506 x 10‘4 4.732 x 10‘4 0.10 4.448 0.4536 0.373 0.138 1.10 x 10"3 9.46 x 10"4 5.03 2.579 x 10"4 4.85 x 10'6 2.59 x 106 14.59 0.229 1.0 x 10"4 1.016 x 103 1.0 x 103 9.072 x 102 1.333 x 102 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-Farrar Cost Indexes {Oil and Gas /.) , 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 All these indexes were developed with various elements, such as material availability and labor productivity, taken into account. However, the pro­ portion 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, 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 TABLE 1

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

Chemical Engineering and Marshall and Swift Plant and Equipment Cost Indexes since 1950 CE Index M&S Index Year CE Index M&S Index 167.9 1971 321.3 73.9 132.3 80.4 180.3 1972 137.2 332.0 144.1 344.1 81.3 180.5 1973 398.4 84.7 182.5 1974 165.4 184.6 1975 182.4 444.3 86.1 472.1 88.3 190.6 1976 192.1 505.4 204.1 93.9 208.8 1977 545.3 98.5 225.1 1978 218.8 599.4 99.7 229.2 1979 238.7 234.5 1980 261.2 659.6 101.8 237.7 721.3 102.0 297.0 1981 101.5 237.2 1982 314.0 745.6 238.5 1983 102.0 316.9 760.8 780.4 102.4 239.2 322.7 1984 1985 325.3 789.6 103.3 241.8 104.2 318.4 797.6 244.9 1986 107.2 252.5 1987 813.6 323.8 262.9 342.5 852.0 109.7 1988 355.4 895.1 273.1 113.6 1989 119.0 285.0 1990 357.6 915.1 125.7 303.3 1991 361.3 930.6 1992

358.2

943.1

xv

xvi

Bringing Costs up to Date

TABLE 2

Date 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 1992

Nelson-Farrar Inflation Refinery Construction Indexes since 1946 (1946 = 100) Nelson-Farrar Materials Labor Miscellaneous Inflation Component Component Equipment Index 100.0 100.0 100.0 100.0 122.4 113.5 114.2 117.0 139.5 128.0 122.1 132.5 143.6 137.1 121.6 139.7 149.5 144.0 126.2 146.2 164.0 152.5 145.0 157.2 164.3 163.1 153.1 163.6 172.4 174.2 158.8 173.5 174.6 183.3 160.7 179.8 176.1 189.6 161.5 184.2 190.4 198.2 180.5 195.3 201.9 208.6 192.1 205.9 204.1 220.4 192.4 213.9 207.8 231.6 196.1 222.1 207.6 241.9 200.0 228.1 207.7 249.4 199.5 232.7 205.9 258.8 198.8 237.6 206.3 268.4 201.4 243.6 209.6 280.5 252.1 206.8 212.0 294.4 211.6 261.4 216.2 310.9 220.9 273.0 219.7 331.3 226.1 286.7 224.1 357.4 304.1 228.8 234.9 391.8 239.3 329.0 250.5 441.1 254.3 364.9 265.2 499.9 268.7 406.0 277.8 545.6 278.0 438.5 292.3 585.2 291.4 468.0 373.3 623.6 361.8 522.7 421.0 678.5 415.9 575.5 445.2 729.4 423.8 615.7 471.3 774.1 438.2 653.0 516.7 824.1 474.1 701.1 573.1 879.0 515.4 756.6 629.2 951.9 578.1 822.8 693.2 1044.2 647.9 903.8 707.6 1154.2 622.8 976.9 712.4 1234.8 656.8 1025.8 735.3 1278.1 665.6 1061.0 739.6 1297.6 673.4 1074.4 730.0 1330.0 684.4 1089.9 748.9 1370.0 703.1 1121.5 802.8 1405.6 732.5 1164.5 829.2 1440.4 769.9 1195.9 832.8 1487.7 797.5 1225.7 832.3 1533.3 827.5 1252.9 824.6 837.6 1579.2 1277.3

xvii

Bringing Costs up to Date

Index. Table 2 shows the Nelson-Farrar Inflation Petroleum Refinery Con­ struction 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/AY

(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 ex­ ponent); CA = actual cost of unit A; and CB = the cost of B being sought for the same time period as cost CA. 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* = CAIB/IA

(2)

Here, CA = old cost; IB = current index value; and IA = index value at the date of old cost. Combining Eqs. (1) and (2), CB = Ca(B/AY(I b/Ia)

(3)

For example, if the total investment cost of plant A was $25,000,000 for 200-million-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), CB

=

C a (B/AY(I b IIa)

= 25.0(300/200)° 6(318.4/165.4) = $61,200,000 JOHN J. McKETTA

Encyclopedia of Chemical Processing and Design 45

Project Progress Management Introduction What can be done to ensure that projects are completed within budget and schedule? The answer depends upon many things. It may not be possible to complete a project within budget and schedule if the estimate is too low and the schedule too optimistic. Also, if the job scope undergoes continuing change throughout project execution, it will be difficult, if not impossible, to control a project effectively. However, if proper control procedures are set up and followed throughout project execution, costs can be controlled and projects completed within schedule (see Table 1). If there are adverse cost and schedule trends, a properly applied project control system will indicate these trends early enough so that decisions can be made to bring the project back in line. The basic requirements of a good project control system are first to prepare a detailed plan for the project, then measure performance against that plan, and note deviations from the plan so corrective action can be taken. Good project control requires discipline.

Major Cost Elements The cost of a project is made up of three major elements: 1. 2. 3.

Equipment and material expenditures (50-60%) Home office salaries and expense (10-20%) Field construction costs (20-30%)

TABLE 1

10 Steps to Project Control

1. Define the overall project work scope into identifiable work packages. 2. Determine the organizational entities that will perform the work. 3. Assign single-point responsibility for each work package to an organizational entity. 4. Schedule the work packages, establishing durations, sequences, and interrela­ tionships. 5. Budget the work packages in cost, labor-hours, and quantity units. 6. Prepare performance measurement baselines by distributing the budgets for the work packages over the schedule time for the work packages. 7. Measure physical accomplishment for each work package in a quantitative and objective manner. 8. Measure actual expenditures incurred for each work package in the same units as budgeted. 9. Compare the accomplishment and the expenditure against the baseline plan periodically throughout the project execution. 10. Report and monitor against the plan. Note deviations and extend, extrapolate, or trend to give at-completion values. Initiate positive action to correct adverse trends.

1

Project Progress Management

Of these, equipment and material expenditures are controlled through ma­ terial management and quantity control procedures, combined with good requisition and purchase order control procedures [1]. Home office and field labor costs can be more difficult to control. These costs are directly related to performance and productivity. A decline in productivity will cause a cost overrun. It will also mean that progress is not being achieved at the planned rate. This, in turn, will lead to schedule slippage. Conversely, a schedule delay for other reasons can contribute to reduced productivity if full efficiency cannot be maintained. In the home office, the key item to control is the hours expended by the major engineering disciplines. In the field, it is the hours expended by the major labor crafts. It is in these areas that effective progress and performance measurement systems play an important role in helping to monitor and control project costs and provide realistic trends and projections of the cost at completion. For effective control, the first requirement is to prepare a plan against which one can measure and compare. The first 6 of the 10 steps to project control all have to do with planning, leading to the preparation of perform­ ance measurement baselines.

Selecting Organizational Entities In the home office, performance measurement baselines are required for each of the major engineering organizational entities or disciplines, such as process, equipment, civil, structural, piping, electrical, and instruments. In the field, it is primarily the same basic organizational entities or craft skills, such as civil, equipment, structural, piping, electrical, and instruments (see Fig. 1).

Work Packages and Work Elements The work scope is subdivided into successively greater detail under the headings of work packages and work elements. A work package is a generic grouping of work elements. Each element has a similar nature, requires a common resource, has the same unit hourly rate, and occurs over a defined period within the overall schedule duration for the organizational entity. A work package is made up of work elements, which have quantities to be installed or numbers of units to be performed, and a unit hourly rate. The estimate of hours for the work package is the sum of the estimate of hours for each work element, which in turn is the sum of the quantities to be installed or performed times the unit hourly rate (see Fig. 2). To illustrate this a little further: in the home office, a work package for the instrument group might generically be instrument data sheets, and a work package for the piping group might be underground piping drawings.

Project Progress Management

3

Helping Helping Helping

Helping

Helping

Helping

Helping

Helping

Helping Helping

Helping

Helping

Helping

FIG. 1

Helping

Helping

Helping

Helping

O rganizational entities requiring perform ance m easurem ent.

The work elements within the instrument work package could be level data sheets, flow data sheets, pressure data sheets, and so on. For each of these types of data sheets, there will be a standard number of hours or hourly rate per data sheet. If the number of data sheets are quantified, then the budget per work element is the number of data sheets times the hourly rate. The budget per work package is the sum of the hours for the work elements.

Project

Section 1

Section 2

Engineering

Materials

Civil

Equipment

Work Pkge 1

Construction

Structural

Work Pkge 2

Work Work Work Work

FIG. 2

Section 3

element element element element

Piping

Work Pkge 3

1 2 3 4

Establishing work packages and work elements per organizational entity.

Project Progress Management

For the field, the situation is much the same. For the civil craft, a generic work package might be concrete foundations. Specific work elements might be large-pour foundations, medium-pour foundations, or small foundations. Each has a different unit hourly rate. Large-pour foundations require less hours per cubic yard than small-pour foundations. For the piping craft in the field, a generic work package would be various types of pipe installation, such as above-ground, below ground, carbon steel, or alloy. Within each work package would be the work elements, being the individual sizes of pipe for which there is a specific unit hourly rate. The budget per work element would be the quantities of pipe to be installed times the unit hourly rate. The overall budget for the work package would be the sum of the budget hours for the work elements.

Linking Work Packages to the Schedule Having established the work packages, the work elements, the quantities, the unit hourly rates, and the budgets, we have quantitatively defined the work. We have not, as yet, related it to the schedule or the time during which the work must be done, that is, the rate at which the work must be completed to meet the project execution plan. The work scope, which has been quantified, must now be linked to the schedule. This is done by listing the work packages and/or the work elements with their quantities and bud­ geted hours. Against each, a bar schedule is inserted indicating the duration for the activity of the work element or the work package. Whether this is done at the work element or work package level depends upon the uniformity of the work elements in a work package and the sched­ uling sequence of the work elements. If, within a work package, there is general uniformity in the work elements, all have the same unit hourly rates, and all occur in the same schedule period, then it is adequate to use a work package budget and schedule to establish a performance measurement base­ line. If, however, work elements are performed at different times in a sched­ ule and there is some variation in the hourly rates or budgets for work elements, then go to the next level of detail and create a bar chart schedule for the major work elements.

Time Phasing Budgets over Schedule Once the work package and work element list has been prepared with the budget and the bar chart schedule, the next step is to distribute or time phase the budget hours over the schedule duration for the work element. This distribution should reflect the manner in which the labor will be applied. It could be a straight-line distribution (an even number of hours expended over each period from start to finish), or it could be an uneven distribution

Project Progress Management

5

reflecting some form of loading curve, indicating a mobilization or buildup to a peak activity and then a decline as the activity is completed and destaffed (see Fig. 3). If the hour distribution is being spread over a large number of work elements, the need to use a loading curve shape for the distribution of hours diminishes. This is because the phasing and sequencing of the work elements themselves will create a bell-shaped loading curve when they are totalled for the overall organizational entity. However, if the budget is distributed in larger amounts at the work package level, then a curve shape for the distribution is more desirable.

Preparing Loading and Progress Curves Once the hours have been distributed to each working period over the schedule activity duration, then the sum of the distributions per period can be totaled to give an overall loading curve for the organizational entity. This loading curve will generally be a bell shape, reflecting the hours spent for each reporting period. Reporting periods may be days, weeks, or months. For most projects, a week is the most appropriate reporting period. For very large extended projects, a month may be acceptable. For short-duration projects, such as turnarounds or retrofits, it may be necessary to use days or even hours for the reporting period.

Work Work Meas pckge elmnt unit

|

Oty

Unit rate

Budget hours

R

(D

£

«

13

CO

®

I I I

Helping

Helping

Helping

(D .tO

Helping

Helping Helping

Helping

Helping Helping

W~

5

FIG. 1

£

3.0 ) 92.2 4.5 0.3 0

Residue Gas Spec 5.0 sum

Design n

CO, c C, C4+

100.0 HHV

TABLE 5

7.8

2.4

89.5 2.5 0 0 0.2

95.6 1.5 0.1 0.2 0.2

100.0 910

100.0 979

2

Q

Mol % December 86

Design

C, CO,

c

2

C,

C4+

3.0

N G L and Recoveries: Design and December 1986 Recoveries (% )

NGL (Liquid vol %)

N-.

Spec

December 86

0 0 0.3 2.5 45.0 52.2

0 0 0.2 25.2 35.4 39.2

100.0

100.0

145.0

83.0

Spec

1.0

Design

Typical

70 99 3-6 95 100

80 99.8 10-35 80-95 100

200 psig bubble point, °F

Propane Recovery, Ryan/Holmes Process

Comparison of Propane Recovery and Ethane Recovery Modes Recovery of C 3 is easier than recovery of C2. C 2 and C 0 2 boil close to one another and form an azeotrope; C 3 and C 0 2 have about a 90°F separation of boiling points but are nearly azeotropic on the C 0 2-rich end [3-5]. To recover C2, the main C 0 2/NGL column (for C 2 recovery, designated the ethane recovery column (ERC)) has to be designed and run for a sharper sep­ aration, requiring more additive, more reflux, more reboil, more trays, and more column cross-sectional area per unit of feed. The basic equilibria and their improvement in the presence of additive are shown in Fig. 6 . The equilibria presented are based on conditions near the top of the main C 0 2/NGL column. The data in Fig. 6 are specifically the values obtained in a single vapor-liquid equilibrium stage using plant compositions and conditions. The data are reasonably representative of the separability of the key pairs, C 0 2 /C 2 and C 0 2 /C3, in the part of the column that performs the main fractionation, the section above the main feed and below the tray where the recycle additive enters. Considering the separation of C 0 2 and C 3 with no additive present, their relative volatility is about 1.50. Small amounts of additive, 5-20% in the liquid phase, cause the relative volatility to be improved to 2.0-4.0, and the separation is made much easier [6 ]. Figure 6 shows that small amounts of additive are effective to help C 3 recovery but ineffective to aid in C2 recovery. When more additive is used, C2 recovery is possible; it also follows that C3 recovery would be very high. With no additive present, or even 5-20% additive, the relative volatility of C 0 2 to C2 is less than 1 . 0 and a substantial separation and recovery of C 2 is impossible. At least 25-35% additive is needed to get into the area where the relative volatility is greater than 1.0 and substantial C 2 recovery becomes possible. Relative volatility a is a key parameter in column design, but 1/In a is more directly reflective of the actual fractionation difficulty. Given com­ parable product purities, feed thermal conditions, and sufficient trays, the column height, column cross-sectional area, and reflux, reboil, and additive are proportional to 1 /ln a. The greater difficulty of separations with a close to 1.0 can be seen by the 1/ln a function. For C3 separation without additive, taking a to be 1.45, 1/ln a is 2.7 and the reflux to feed ratio is about 2.7 with plenty of trays. Raising a to 3.1 with 13% additive lowers 1/ln a to 0.9 and the separation is three times easier, needing one-third the reflux, reboil, trays, and area. Figure 7 plots the relative volatility data of Fig. 6 as 1/ln a. The same conclusions can be drawn as before, but the practical area of operation can be anticipated from the “shape” of the curves. For C 3 recovery it appears 5-25% additive ought to be the best range. For C2 recovery 40% or more additive would seem to be practical. The comparison of C2 and C 3 recovery modes involves more than the main C 0 2/NGL column. The A R C is considerably different in the two cases. The additive flow rate is larger for C2 recovery, which directly affects the

Propane Recovery, Ryan/Holmes Process

73

Helping

Helping

HelpingHelping

% C5+ additive in liquid phase FIG. 6

COVC2 and C 02/C3 vapor-liquid equilibrium improvement with additive.

Propane Recovery, Ryan/Holmes Process

Helping

Helping

% C 5+ additive in liquid phase

FIG. 7

C 0 2/C2 and C 0 2/C3 separation difficulty.

Propane Recovery, Ryan/Holmes Process

75

ARC cross-sectional area and the reboiler size and duty. For the C 2 recovery case, the overhead will have a relatively high C 2 content, and it is practical to use a partial condenser for reflux with a vapor net product to utilize ambient cooling. For the PRC case, the content of C 2 is lower and it is likely that the overhead can be totally condensed. As regards the recycle additive composition, a key parameter is to strip out C3. C 3 in the additive enters the main column at a point from which it can slip into the overhead and be coproduced with the C 0 2, especially in the case of the propane recovery mode since less reflux is used, which thus lets a bigger proportion of C3 slip to the overhead product. Horsepower and fuel consumption are key operating cost parameters. The comparison of C2 and C3 recovery can be stated semiquantitatively based on these utilities. Refrigeration horsepower for C2 recovery is two to three times the C3 recovery mode. In the C3 recovery mode, the refrigeration horsepower is a relatively small fraction of the horsepower for the gas compression and reinjection. The reboil fuel requirement of the C 2 recovery mode is about three to four times the C3 recovery mode. The Alvord South plant has shown that there is truly a continuous tran­ sition from C 3 recovery to C2 recovery. The plant was originally designed for C 3 recovery. Of the C2 from the plant feed, 97% was contained in the C 0 2 reinjectant stream; about 3% was recovered in the liquids product. Column conditions were chosen to obtain 90% or more C3 recovery. In the recent past, the plant inlet gas rate has been about 4.3 MMscfd. Capacity was designed at 7.5 MMscfd; actual compared to design is 57%. Mitchell can use the entire capacity of refrigeration, additive recirculation, and heat input to intensify the fractionation; 15-35% C 2 recovery has been obtained, giving the plant credit for liquids production considerably beyond the original design, which planned for only 2-4% C2 recovery. With that much C 2 in the liquids, the temperature must be less than 80°F not to exceed the actual vapor pressure criterion. Detailed analysis of the main column fractionation shows that all the C3 in the overhead (more than 99.9%) comes from C 3 in the additive. Much less than 0.1% of the C 3 in the feed passes directly to the column overhead.

Conclusion The Mitchell Alvord South C 0 2 recovery plant is performing well, and operation is relatively stable. The Ryan/Holmes process satisfied the design criteria of low cost and fast schedule by allowing use of existing equipment within the Mitchell organization. An added bonus was being able to meet the design criteria while having the lowest operating cost. C 0 2 and liquids recovery have been high and satisfactory relative to project design. Given the variations in inlet gas rate and composition, Mitchell believes the Ryan/ Holmes process has performed well and was the correct choice for the application.

76

Propane Refrigeration Systems

Notation Ci, C2, . . . K

alkanes CH4, C2 H6, . . . , per gas production convention vapor-liquid equilibrium constant; K is defined as yjxh where yt is real fraction of component i in the vapor xt is real fraction of component i in the liquid relative volatility (vapor-liquid equilibrium); a is defined as Kf/Kj for component i relative to component j

Much of this article was excerpted with special permission from Energy Progress, December 1987, copyright © 1987 by American Institute of Chem­ ical Engineers, New York, New York 10017.

References 1.

F. G.

Russell,

“ Applications of the DELSEP Membrane System,”

Chemical

Engineering Progress, 80, 48 (October 1984). 2.

J. M. Ryan and J. V. O ’Brien, “ Distillation Technology Increases Propane Recovery in Carbon Dioxide Floods,” Oil and Gas Journal, 62 (October 6, 1986).

3.

F. H. Poettman and D. L. Katz, “Phase Behavior of Binary Carbon DioxideParaffin Systems,” Industrial Engineering Chemistry, 37(9), 847 (September 1945). W. W. Akers, R. E. Kelley, and T. G. Lipscomb, “ Carbon Dioxide-Propane System,” Industrial Engineering Chemistry , 46(12), 2535 (December 1954). H. H. Reamer, B. H. Sage, and W. N. Lacey, “Phase Equilibrium Hydrocarbon Systems,” Industrial Engineering Chemistry , 43, 2515 (1951). F. W. Schaffert and J. M. Ryan, “ Ryan/Holmes Technology Lands E O R Proj­ ects,” Oil and Gas Journal, 133 (January 28, 1985).

4. 5. 6.

JAMES M. RYAN PAT McCANN JOHN V. O ’BRIEN

Propane Refrigeration Systems Introduction The requirements for horsepower and condenser duty can be estimated from generalized charts for propane refrigeration systems.

Propane Refrigeration Systems

77

The ultimate heat sinks for propane refrigeration systems will be cooling water and ambient air. The charts in this article apply over a wide range of condensing temperatures for the refrigerant. Hence, we can use these charts with cold water as the heat sink or without ambient air where cooling water may not be available. In developing the charts, the following assumptions were made: Polytropic efficiency of 0.77 Equal compression ratios between stages Pressure drop of 1.5 psi between the evaporators or economizers and the compressor suction and side-load inlet nozzles Pressure drop of 10 psi across the propane refrigerant condenser The vapor pressure curve for propane refrigerant is presented in Fig. 1, from which pressures and temperatures within the propane refrigeration systems can be developed. The compression ratio r can be calculated from

( 1) where Ps = suction pressure, psia; Pd = discharge pressure, psia; and N = number of stages of compression.

Single-Stage Systems The gas horsepower (GHP) required per million Btu/h (MMBtu/h) of re­ frigeration duty for a single-stage propane refrigeration system at various evaporator and refrigerant condensing temperatures can be found from Fig. 2. The corresponding condenser duty can be determined from Fig. 3. Example 1 Let us estimate the horsepower and condenser duty require­ ments for a single-stage propane refrigeration system (as shown schemati­ cally in Fig. 2) that provides 30 MMBtu/h of process chilling at - 10°F. The refrigerant can be condensed at 100°F. For a condensing temperature of 100°F and an evaporator temperature of - 10°F, we find that the unit (GHP) from Fig. 2 is 187.5 hp/(106 Btu)(h) of refrigeration duty. And from Fig. 3, we find that the corresponding condenser duty factor is 1.477 MMBtu/h(106 Btu)(h). Hence, the total power and the total condenser duty Qcd are

(GHP) = 30(187.5) = 5,625 hp Q cd = 30(1.477) - 44.31 MMBtu/h

Propane Refrigeration Systems

Temperature,°F Helping + 40

+60

+ 80

+1 0 0

+ 120

Helping Helping Helping

Temperature, °F FIG. 1

Vapor-pressure curve for propane.

+ 1 4 0 + 15 0

Propane Refrigeration Systems

Gas horsepower, h p / ( 1 0 6 Btu)(h) refrigeration

du ty

79

Evaporator tem perature, °F

FIG. 2

Gas horsepower for single-stage propane refrigeration system.

Propane Refrigeration Systems

Total condenser duty, MM

B tu /h /(1 0 6 Btu)(h) refrigeration duty

80

Evaporator temperature,°F

FIG. 3

Condenser duty for single-stage propane refrigeration system.

Two-Stage Systems The gas horsepower and condensing duties for a simple two-stage propane refrigeration system at various evaporator and refrigerant condensing tem­ peratures are shown in Fig. 4 and 5. These charts represent only one load at the lower level; the second stage is essentially an economizer level that is established by assuming equal compression ratios between stages. Example 2 For a simple two-stage propane refrigeration system (as shown in Fig. 4), let us estimate the horsepower and condenser duty to provide 35 MMBtu/h of process chilling at -40°F. Let us also determine the temper­ ature and pressure in the economizer. Propane will be condensed by an air cooler at 120°F. We find the power requirement for the given conditions per million Btu/h of refrigeration duty from Fig. 4 and the unit condenser duty for the same conditions from Fig. 5. Therefore, the total power and total condenser duty for this system become

81

Gas horsepower, h p /(1 0 6 Btu)(h) refrigeration

du ty

Propane Refrigeration Systems

Evaporator tem perature, °F

FIG. 4

Gas horsepower for two-stage propane refrigeration system.

(GHP) = 35(279.0) = 9,765 hp

Qcd = 35(1.71) = 59.85 MMBtu/h To find the economizer pressure, we have to calculate the compression ratio, for which we need the suction and discharge pressures at the nozzles of the centrifugal compressor. We begin by finding that the dew point pres­ sure at the evaporator from Fig. 1 is 15.68 psia at -40°F. Also from Fig. 1, we find that the condenser pressure is 242.19 psia at 120°F. Since we assumed a pressure drop of 1.5 psi at the compressor suction nozzle and a 1 0 psi pressure drop across the refrigerant condenser, we find that

Propane Refrigeration Systems

Total condenser duty, MM

B t u /h /(1 0 6 Btu)(h) refrigeration

duty

82

Evaporator tem perature, °F

FIG. 5

Condenser duty for two-stage refrigeration system.

Ps = 15.68 - 1.5 = 14.18 psia Pd = 242.19 + 10 = 252.19 psia By substituting these pressures into Eq. (1), we compute the compression ratio per stage for this system as r = (252.19/14.18)1/2 = 4.217. From this ratio, we calculate the interstage pressure as 14.18(4.217) = 59.79 psia and the economizer pressure as 59.79 + 1.5 = 61.29 psia. With this last pressure, we find that the economizer temperature from Fig. 1 is 25°F. Example 3 Let us estimate the horsepower and condenser duty for the twostage propane system shown in Fig. 6 . On examining the data in Fig. 6 , we find that the first-stage evaporator handles a chilling load of 25 MMBtu/h and the second stage, a load of 10 MMBtu/h, with a subcooler providing additional chilling of the propane refrigerant amounting to 4 MMBtu/h. We will consider the chilling load of 25 MMBtu/h at -40°F indepen­ dently, as a simple two-stage system, which is exactly similar to that of Example 2. We will combine the second refrigeration load of 10 MMBtu/h at 25°F and the refrigerant subcooler duty of 4 MMBtu/h as a single-stage system. From Fig. 2, we find that the (GHP) per 106 Btu/h at an evaporator temperature of 25°F and a refrigerant condensing temperature of 120°F is 144 hp/(106 Btu)(h), and from Fig. 3, the condenser duty at the same tem-

Propane Refrigeration Systems

83

Helping Helping

Helping

FIG. 6

Helping

Two-stage propane system for Example 3.

peratures is 1.366 MMBtu/h/(106 Btu)(h). Hence, for this example, the total horsepower and condenser duty are obtained as the sums for each stage of the system:

(GHP) = 25(279.0) + (10 - 4)(144.0) = 7,839 hp

Qcd = 25(1.71)

4-

(10 - 4)(1.366)

= 50.95 MMBtu/h

Three-Stage Systems The gas horsepower and condenser duty for a simple three-stage propane refrigeration system at various evaporator and refrigerant condensing tem­ peratures are shown in Figs. 7 and 8 . A simple system (as shown in Fig. 7) consists of one refrigeration load at the lower level and two interstage econ­ omizers. Temperature and pressure conditions in the economizers are es­ tablished by assuming equal compression ratios between the stages. Example 4 Let us determine the refrigeration requirements for a simple three-stage propane system designed to provide 40 MMBtu/h of process chilling at - 40°F. Let us also determine the interstage economizer conditions for this system. We will use cooling water at 8 8 °F to condense propane at 100°F. From Figs. 7 and 8 , we find that the (GHP) and condenser duty factors at an evaporator temperature of - 40°F and a refrigerant condensing tem­ perature of 100°F are 224.2 hp and 1.5705 MMBtu/h per 106 Btu/h of refrigeration duty, respectively. Therefore, total horsepower and condenser duty become

o

11

-4 0

h

11

rn

ir r r r m r n

0

20

-2 0

r - n r r r m T m 11 40

m

60

h

p s ti

80

rw

i i rk i

100

r u

jj r u

i i r u

120

i i t h

140

Evaporator tem perature, °F

FIG. 7

-4 0

-2 0

0

Gas horsepower for three-stage propane refrigeration system.

20

40

60

80

100

120

Evaporator tem perature, °F

FIG. 8

Condenser duty for three-stage propane refrigeration system.

140

Propane Refrigeration Systems

85

(GHP) = 40(224.2) = 8,968 hp

Qcd = 40(1.5705) = 62.82 MMBtu/h By using Fig. 1, we obtain the dew point pressure at -40°F in the evaporator as 15.68 psia and the condenser pressure at 100°F as 188.32 psia. Subtracting or adding the appropriate pressure drops at the compressor nozzles, we find that Ps = 14.18 psia and Pd = 198.32 psia. Substituting these pressures into Eq. (1) yields the compression ratio r = (198.32/14.18),/3 = 2.4093. Using this ratio, we can calculate the firststage PJ(l) and the second-stage PJ{2) discharge pressures for this system:

Pd(l) = 2.4093(14.18) - 34.16 psia Pd(2) = 2.4093(34.16) - 82.31 psia To the first-stage discharge pressure of 34.16 psia, we add 1.5 psi for pressure drop to get 35.66 psia as the second-stage economizer pressure, and from Fig. 1, we find that the second-stage economizer temperature is -3.9°F. Similarly, the third-stage economizer pressure becomes 82.31 + 1.5 = 83.81 psia and the temperature for the third-stage economizer is 43.9°F.

Helping Other Three-Stage Systems Examples 5 and 6 illustrate the use of these charts for some possible com­ binations of three-stage systems. Example 5 For the conditions developed in Example 4, let us determine the horsepower and condenser duty requirements for the three-stage pro­ pane system whose chilling loads and subcooling duty are shown in Fig. 9.

Helping

Helping

Helping

Helping

Helping

Helping FIG. 9

Three-stage propane system for Example 5.

Propane Refrigeration Systems We will consider the chilling loads as follows: first-stage load of 23 MMBtu/h as a simple three-stage system; second-stage load of 10 MMBtu/h as a simple two-stage system; third-stage load of 7 MMBtu/h to be combined with subcooling load of 3 MMBtu/h to yield a net chilling load of 4 MMBtu/h, acting as a simple single-stage system. Let us summarize the analysis and solution for this problem: Load, M M B tu/h Load temperature, °F Condenser temperature, °F Stages, number (G HP) factor, hp/(106 Btu)(h) Condenser duty factor, M M Btu/h/(106 Btu)(h)

23 -40 100 3 Fig. 7 224.2 Fig. 8 1.5705

10 -3.9 100 2 Fig. 4 148.0 Fig. 5 1.3766

4 43.9 100 1 Fig. 2 74.5 Fig. 3 1.190

Total compression horsepower and total condensing duty are obtained by adding the amounts for each load: (GHP) = 23(224.2) + 10(148.0) + (7 - 3)(74.5) (GHP) = 6,935 hp

Qcd = 23(1.5705) + 10(1.3766) + (7 - 3)(1.190) Qcd = 54.65 MMBtu/h The next example will illustrate the use of these charts for the cascading of refrigerant systems. Here, we will consider the cascading of a two-stage ethane system into the first stage of a three-stage propane system. For a summary on cascaded refrigeration systems, see the article on propylene refrigeration systems in this encyclopedia. Example 6 The two-stage ethane refrigeration system is taken from Ex­ ample 3 in the article on ethane refrigeration in this encyclopedia and cas­ caded into the three-stage propane system of the previous example. The schematics for each of these systems and the accompanying data are shown in Fig. 10. Let us estimate the total condenser duty for the propane system and the total horsepower required for the cascaded system. The horsepower and condenser duty for the ethane refrigeration system (Fig. 10) are 3,424 hp and 30.71 MMBtu/h, respectively, as determined in the article on ethane refrigeration in this encyclopedia. Since this condenser duty is cascaded into the first-stage suction drum of the propane system, the total propane refrigeration duty at -40°F becomes 23 + 30.71 = 53.71 MMBtu/h. We will use the horsepower and condenser duty factors from Example 5 to calculate the requirements for the three-stage propane system that now handles the additional condenser duty from the ethane system. The results are as follows:

87

Propane Refrigeration Systems

Helping Helping Helping Helping Helping

FIG. 10

-4 0

-2 0

Helping HelpingHelping

Cascaded ethane to propane system for Example 6.

0

FIG. 11

20

40 60 Temperature, °F

80

100

Enthalpies of propane for liquid and vapor.

120

140

88

Propylene Oxide and Glycol

(GHP) = 53.71(224.2) + 10(148.0) + (7 - 3)(74.5) (GHP) = 13,820 hp

Qcd = 53.71(1.5705) + 10(1.3766) + (7 - 3)(1.190) Qcd = 102.88 MMBtu/h The compression horsepower required for the cascaded ethane-propane refrigeration system is as follows: System

(GHP)

Ethane Propane Total

3,424 hp 13,820 hp 17,244 hp

Liquid and vapor enthalpies for propane, applicable over the refriger­ ation temperature range, are presented in Fig. 11. We may use these enthalpy data to determine refrigerant flow rates through the system, as demonstrated in Example 7 of the ethane refrigeration article in this encyclopedia. Much of this article was excerpted with special permission from Chemical

Engineering, March 26, 1979. YUV R. MEHRA

Propylene Oxide and Glycol Introduction Propylene oxide (PO) is an important propylene-derived chemical. In the United States, it is estimated that PO is the third largest derivative of pro­ pylene, consuming over 11% of propylene supplied in 1985 [1]. It is the starting material for polyether polyols, which is one of the major reactants in the preparation of urethane-based materials. There are two basic types of processes that have been significant in the commercial manufacture of propylene oxide. The first and oldest is the classic chlorohydrin process, which still accounts for the majority of the world PO capacity. The second is the newer and more economic direct oxidation route, whose development and commercialization in the 1960s spurred the increased use of urethane chemicals. The largest portion of new capacity built in the western world since the late 1960s has been based upon the direct oxidation route.

Propylene Oxide and Glycol

89

Plant Capacities The total propylene oxide capacity for the western world (excluding Eastern Europe and the People’s Republic of China) is estimated at 2.9 million t, as shown in Table 1. Capacities for Eastern Europe and the People’s Re­ public of China are estimated at about one-tenth that of the western world, or about 300,000 t. Announced capacity additions will push capacity for PO by 1991 to 3.1 million t in the western world. As discussed in detail in Process Description, there are currently three processes used for the manufacture of PO: (1) chlorohydrin; (2) direct ox­ idation with tertiary butyl alcohol (TBA) as a coproduct; and (3) direct oxidation with styrene monomer (SM) as a coproduct. Table 2 lists the

TABLE 1 Regional Propylene Oxide Capacity, 1986a Western Hemisphere Western Europe Asia Total

Thousand t

%

1,515 1,090 275 2,880

53 38 9 100

'Industry sources.

TABLE 2

Worldwide Propylene Oxide Capacity, 1986

Company

Country

Thousand t

A R C O Chemical A R C O Chemical Dow Chemical Dow Chemical Dow Quimica Atochem A R C O Chemie Nederland Shell Nederland Chemie BASF Dow Chemical Erdoelchemie Montedison Petrochemica Enpetrol Nihon Oxirane Asahi Glass Mitsui Toatsu Showa Denko Tokuyama Soda Chiunglong Petrochemical

USA USA USA Canada Brazil France Netherlands Netherlands West Germany West Germany West Germany Italy Spain Japan Japan Japan Japan Japan Taiwan

240 476 635 64 100 65 230 125 65 380 140 40 45 100 55 36 33 36 15

aNot operating in 1987.

Technology PO/SM PO/TBA Chlorohydrin Chlorohydrin Chlorohydrin Chlorohydrina PO /T BA PO/SM Chlorohydrin Chlorohydrin Chlorohydrin Chlorohydrin PO/SM PO/SM Chlorohydrin Chlorohydrin Chlorohydrin Chlorohydrin Chlorohydrin

Propylene Oxide and Glycol

current capacities by region and by company, indicating which of the pro­ cesses are used. In 1986, 58% of the world’s capacity was based upon chlo­ rohydrin technology, 24% based upon PO/TBA, and 18% upon PO/SM. By 1991, it is expected that almost 50% of the world PO capacity will be based upon the newer direct oxidation technology. This is a continuance of a trend begun in the late 1960s with the construction of the first direct oxidation plant. Table 3 shows how the proportion of a PO capacity based upon direct oxidation has grown in the western world.

Consumption It is estimated that in 1986, 2.4 million t propylene oxide was consumed in the world. As shown in Table 4, the USA and Western Europe consumed 67% of that amount. However, because of the large international trade in propylene oxide, actual production of propylene oxide in the USA and Western Europe amounted to 82% of worldwide production. Table 5 shows the movements of propylene oxide between the regions. Since it is easier and more economical to ship the derivatives of propylene oxide than it is to ship propylene oxide itself, most of the world trade in propylene oxide is done in the form of derivatives. Thus, although Table 5 shows the production and consumption of propylene oxide, most of the propylene oxide that moves from region to region is in the form of deriv­ atives. Table 5 shows the estimated amount of propylene oxide contained in those derivatives, not the actual quantity of the derivatives themselves. The major uses for propylene oxide are in the manufacture of polyether polyols and propylene glycol (Table 6 ). In 1986, almost 90% of the PO produced was consumed in these two uses. Other uses for propylene oxide

TABLE 3

Propylene Oxide Capacity by Type (% )a

Chlorohydrin Direct oxidation PO /T B A PO/SM

1965

1975

1985

100

74

58

0 0

24 2

25 17

'Industry sources.

TABLE 4

Regional Propylene Oxide Demand, 1986a Thousand t

Western Hemisphere Western Europe Asia Total 'Industry sources.

1,053 905 407 2,365

% 45 38 17 100

Propylene Oxide and Glycol

TABLE 5

91

International Propylene Oxide Balance, 1986 (Thousand t):l From

To

Rest of Western Hemisphere

USA

USA Rest of Western Hemisphere Western Europe Eastern Europe Africa and Mideast Asia Total production

Western Europe

802

12

7

71 16

151

9 746 84 59 50 955

—

—

—

—

—

Ill 1,000

—

163

Asia —

— — — —

247 247

Total Consumption 821 231 762 84h 59 408 2,365

'PO and equivalent PO in derivatives from industry estimates. 'Imports: not including domestic production.

include glycol ethers and acetates, functional fluids, and a variety of specialty compounds. Polyether polyols, which consumed 63% of worldwide propylene oxide production, is a family of materials made by reacting propylene oxide with a compound containing two or more active hydrogens. Propylene oxide is added until the desired molecular weight is achieved. The most widely used polyether polyols are of 3,000 and 3,500 molecular weight. When combined with an isocyanate, either toluene diisocyanate or PMDI, and suitable cat­ alysts and blowing agents, a polyurethane material is produced. Propylene glycol is the second most important use for propylene oxide. When propylene oxide is reacted with water, propylene glycol is produced, along with its oligomers, such as dipropylene glycol and tripropylene glycol. The largest market for propylene glycol is the manufacture of unsaturated polyester resins. The other major uses that take advantage of the humectant property of propylene glycol are in the manufacture of cosmetics, tobacco, and pet foods. One of the fastest growing uses for propylene oxide is in the manufacture of propylene glycol ethers. The U.S. Environmental Protection Agency (EPA) has found that ethylene glycol ethers cause birth defects in animals

TABLE 6 Worldwide Propylene Oxide Uses, 1986 (Thousand t)a Polyols for Flexible and semirigid foam 1,170 Rigid foam 310 Propylene glycols 568 Propylene glycol ethers 67 Miscellaneous 250 Total__________________________________ 2,365 'Industry sources.

Propylene Oxide and Glycol

92

and pose health hazards to humans. Thus, the less toxic propylene glycol ethers are being used as substitutes for the more toxic ethylene glycol ethers in paints and solvents.

Process Overview There is a long history of research efforts to produce propylene oxide by direct oxidation of propylene with air or oxygen. These efforts are explored here, but none give close to commercial economics. Both commercial routes and several of the unexploited technologies generate an oxidant in a separate step and result in a large volume of coproduct. These are chlorine as oxidant and NaOH by-product for the chlorohydrin route and a hydroperoxide ox­ idant and an alcohol by-product for the oxirane route. Discussions of the economics and technology inevitably require exploration of the economics and technology of producing the coproducts and the markets for them.

Oxirane Process History In the early 1960s there was ample incentive for research on a new propylene oxide process; the market was growing rapidly because of the penetration of urethane foams into cushioning and insulation markets, and the conven­ tional chlorohydrin process had high raw material costs and effluent prob­ lems. Two separate lines of research by Halcon International and Atlantic Richfield (ARCO) moved in the same direction and produced a marriage that resulted in the oxirane process. A R C O (see Example II in Ref. [2a]) explored the oxidation of propylene in inert solvents and found that oxidation catalyzed by Mo, Ti, W, and V gave significantly higher yields of propylene oxide than reactions catalyzed by Co, Mn, or Fe, though not yet at commercial yields. The poor yields from most catalyzed oxidations result from the following free radical oxi­ dation mechanism reactions [3]. c h 2 = c h - c h 2 o 2- + c h 2 = c h - c h 3— > CH 2 -CH-CH2 0 2H + CH 2 = CH-CH2- (1) h 2 = c h - c h 2- + o 2 — > c h 2 = c h - c h 2 o 2*

(2 )

c h 2 = c h - c h 2 o 2- + c h 2 = c h - c h 3— ► c h 2 = c h - c h 2 o- + PO

CH 2 = CH-CH 2 0 ‘ + c h 2 = c h - c h 3— > CH 2 = CH-CH2OH + CH 2 = CH 2 -CH2* (4)

Propylene Oxide and Glycol

93

R O O H + M "-- > RO- + OH- + M”+1 R O O H + M'!+1 -- > R O / + M/! + H +

(5)

These researchers then decided to produce the peroxide exogenously by /C4 oxidation [2 ]. Since early in the century, it had been known that peracids would oxidize olefins to epoxides [4], and an industry grew up around producing specialty epoxides and epoxidized soybean oil [5]. However, the cost and safe handling of very large amounts of peracetic acid were barriers to using it for making PO. Brill [6 ] showed that hydroperoxides made epoxides in low yield. Halcon set out to catalyze this reaction and found that the same group of transition metal species was significantly more effective than other metals or metalfree systems [7]. Halcon and A RCO formed a joint venture and developed processes to exploit this technology. A process using r-butyl hydroperoxide (TBHP) as oxidant and coproducing TBA and isobutylene was commercialized in the USA and the Netherlands. A second process using ethylbenzene oxidation and coproducing styrene has also been commercialized in Spain, Japan, and the USA. Both are discussed more fully. An alternative catalyst system was developed by Shell [8 ]. These are transition metal oxides supported on main group oxides, with T i0 2 on Si0 2 [9] their preferred combination. It may have certain advantages over soluble molybdenum (Mo) of not requiring continuous makeup or disposal and providing better selectivities with difficult to epoxidize olefins but may have limitations on peroxide conversion. Shell has commercialized a PO plant in the Netherlands with styrene as coproduct.

Chemistry of Epoxidation The facile uncatalyzed epoxidation using the peracids is likely due to the acid nature of the leaving group and the five-numbered ring transition com­ plex [2a]. Effective metals provide both the acidity and a useful transition state. The preferred mechanism is [10, 11]

Helping

Helping

Helping

Propylene Oxide and Glycol

However, other similar mechanisms [12] cannot be ruled out. Many molybdenum compounds, such as molybdenum trioxide, molybdenum naphthenate, molybdenum hexacarbonyl, and molybdenum metal, are readily converted to active catalyst [13]. The molybdenum is rapidly oxidized to its most stable valence, 6 , and takes on bidentate glycol ligands [13]:

Helping

Helping

One can stop the epoxidation by complexing the molybdenum with very strong ligands [14] or by adding base, which will precipitate the molybdate salt. Alcohols and water will inhibit the reaction [12] because of competition for the available site on the molybdenum: H / O O

Helping — Mo + ROH

/I

Helping

/ I \

(9)

where R = H, alkyl, or alkaryl. According to Sheldon [11], the major nonselective products arise from decomposition of the Mo-ROOH complex in parallel with the olefin reac­ tion: OH I/ — Mo

/ 1I \ O /

/ OR --- > — M o = 0 + RO + OH"

/I1

(10)

A variety of undesirable products then arise from molybdenum reoxidation and the reaction of the alkoxy species. This reaction limits the commercial temperature of operation and is controlled by use of excess olefin.

C4 Coproduct Process About a quarter of the total installed propylene oxide capacity uses isobutane oxidation and makes by-product of f-butyl alcohol. The plants are located in the USA and Europe. Some of the TBA is dehydrated to make highpurity isobutylene, primarily for butyl rubber but also for other isobutylene polymers, alkylation of aromatics to make inhibitors and antioxidants, and other chemical uses. The majority is used as a high octane blending com­

95

Propylene Oxide and Glycol

ponent, frequently in conjunction with methanol. This use is giving way to production of MTBE, also for gasoline blending, via dehydration of r-butyl alcohol to isobutylene. The theoretical weight ratio TBA/PO is 1.32. However, commercial plants produce 2-3 lb TBA per lb PO, depending on demand. The source of and control of the production of the additional TBA is discussed in Oxidation. A description of the remainder of the process follows. Detailed flow sheets have not been provided because of the proprietary nature of the technology. However, subscription services have provided detailed designs with capital estimates and elements of production [15, 16].

Isobutane Oxidation Isobutane oxidation is carried out in the liquid phase at elevated tempera­ tures (130-160°C) and pressures [17, 18]. This is above the critical temper­ ature of isobutane (134°C), so products (TBA and TBHP) must be present to maintain a liquid phase. All commercial installations use oxygen rather than air. It is not obvious that it is easier to separate 0 2 from N 2 in an air plant than isobutane from N 2 vented from the oxidation system. The decision appears to rest with the air separations companies’ pricing philosophy and economics of scale rather than any economics one could generate based on a stand-alone 0 2 plant. A plant using 0 2 produces an off-gas with sufficient heating value to burn economically in a furnace, unlike an air-based plant. This is an advantage that may grow with time as environmental legislation requires greater and greater cleanup of effluents. However, the hazards of mixing pure oxygen with hydrocarbons leads to capital additions (e.g., barricades) and the threat of business interruption. The isobutane oxidation proceeds by the following classic steps [19] R O O H -- > R O ’ + O H ’ RO*

or

*OH + R H -- > R* + R O H

(11) or

H20

(12)

R- + 0 2 -- > R 0 2-

(13)

R 0 2* + R H -- > R 0 2H (TBHP) + R ‘

(14)

2 R 0 2*-- > R O O R (DTBP) + 0 2

(15)

where R = (CH 3 )3 C. However, the unusually large alcohol make, as high as equimolar, leads one to look for additional steps. These are accepted to be [2 0 ] 2 R 0 2*-- > 2RO- + 0 2

(16)

RO* + R H -- ► R O H + R*

(17)

Propylene Oxide and Glycol

In addition to the make of TBA, water, and di-r-butyl peroxide (DTBP) described earlier, other nonselective reactions occur: R 0 2--- > (CH 3 ) 2 CO + CH 3 0-

(18)

The methoxyl radical leads to a range of Q products: methanol, formal­ dehyde, formic acid, formate esters, CO, and C 0 2. The alcohol- and ketone-producing reactions (which would be very severe for the phenol producers) are not particularly detrimental for several rea­ sons: 1.

2. 3.

Because of the low molecular weight of isobutane and the very large gasoline pool that the coproducts (TBA and MTBE) go into, the mar­ keting of 2-3 lb TBA per lb PO is not onerous. The acetone can be recovered. The residual Q products are burned for fuel.

When one considers the relative prices of isobutane, acetone, gasoline, and fuel, it appears that cleavage is not detrimental to the economics. The major drawback is that the by-products— methanol, formic acid, water, and ace­ tone— are detrimental to the epoxidation reaction. Inevitably, abstraction of one of the methyl hydrogens also occurs [18]: R O / + iC4-- > R 0 2H + (CH 3 )2 -CH-CH2-

(19)

The entire range of isobutyl compounds, hydroperoxide, alcohol, aldehyde, and acid result. Increased temperature promotes cleavage and methyl hy­ drogen extraction relative to the desired reactions. Commercial isobutane is 95 wt% minimum purity, the major impurities being propane and n-butane. These are undesirable for the following rea­ sons: 1. 2. 3. 4.

They oxidize more slowly than isobutane. Buildup of propane increases total reactor pressure. Both make undesired products, such as isopropanol, 5 -butanol, and methyl ethyl ketone. The secondary hydrocarbons inhibit oxidation of the tertiary hydrocar­ bons [2 1 ].

Rather than accepting these in full measure, commercial plants alleviate the situation by partial removal of impurities from feed and recycle isobutane. Butanes are available from both fresh production (lights in crude oil and natural gas liquids, NGL) and from refinery processing, for example, cat­ alytic cracking. Some purified isobutane is available from both sources, but the total is insufficient to satisfy A RCO and refinery alkylation demand. Hence, in both Europe and the USA, isomerization of rc-butane [22] is used.

97

Propylene Oxide and Glycol

From this source, the iC4 feed to oxidation will usually be purer than com­ mercial specification because recovered nC4can be recycled to isomerization. Refinery butanes or isobutane can have up to 1% olefins by specification and are commingled with “ natural” C4 ’s, particularly in Europe. These are undesirable because isobutylene in particular is a poison to isomerization catalysts and because allyl oxidations of olefins lead to a series of nonuseful products [reactions (1) through (4)], reduce oxidation rate, and increase acid content. Uncontrolled heavy metals and sulfur are also highly detrimental. They catalyze hydroperoxide breakdown [reaction (5)], with attendant pressure and temperature excursions. Much literature exists on the use of such agents to increase conversion rates. However, the authors are not aware of the use of such materials in commercial oxidations. It appears that comparable rate increases can be obtained simply by raising temperature, and at the same rate, fewer by-products are obtained without catalysts. There is also considerable literature on the use of bases, buffers, and chelating agents [23] to improve yields. Some of these effects appear to be artifacts of high surface area pilot-plant equipment, and some are real. Off-gases from the oxidizer are treated to recycle contained isobutane 1. 2. 3.

Recycles unconverted iC4 Rejects unconverted nC4 and light oxidation products, such as water and acetone Produces a concentrated oxidate suitable for epoxidation

These separations are complicated by the necessity to condense isobutane under pressure and the thermal instability of r-butyl hydroperoxide. This generally leads to two or more columns staged in pressure.

Epoxidation with f-Butyl Hydroperoxide Commercial epoxidation of propylene with TBHP generally operates at 100130°C [24] using 10-300 ppm Mo [24, 25]. The temperatures are generally higher and catalyst concentration lower than academic studies because of the expensive reactor and catalyst. Also, propylene is one of the less reactive olefins [13]. This means that a high propylene concentration is needed to compete with by-product reactions. The high propylene concentration leads to very high pressures and high recycle costs. Propylene is held in the liquid phase at temperatures above its critical point (92°C) by the presence of the oxidate. Selection of process conditions requires a multiparameter optimi­ zation supported by an extensive reaction data set formulated into a kinetic model and also a sophisticated thermodynamic fluid model. Loss reactions fall into two categories, Mo catalyzed and homogeneous. The RO* produced by reaction (10) can either cleave: RO*-- > (CH3)C = 0 + CH3* or abstract an allyl hydrogen from propylene:

(20)

Propylene Oxide and Glycol

RO* (or CHO + CH3-CH = CH 2 -- > •CH2-CH = CH 2 + R O H

(or CH4)

(21)

The allyl radical eventually produces heavy compounds. Propylene oxide can also react with any of the numerous protic species, alcohols, TBHP, water, or acids to produce losses. The chief impurities in chemical-grade propylene are ethane and pro­ pane. These are inert diluents but quite costly to handle because of the already high pressure of the system. They must be separated from fresh and recycle propylene. Polymer-grade propylene virtually eliminates this prob­ lem.

Separations One must separate epoxidation effluent into propylene for recycle, PO and TBA products, and various wastestreams. Specifications for PO and for gasoline-grade r-butyl alcohol (GTBA) are given in Tables 1 and 2. PO largely goes into preparation of polyurethane polyols. Numerous impurities made in oxidation and epoxidation must be removed to exacting levels. Among these are water, methanol, acetaldehyde, methyl formate, C 5 -C7 hydrocarbons, propionaldehyde, and acetone. A combination of di­ rect and extractive distillation is used to remove these impurities [26]. Treatment of TBA depends on its ultimate use. GTBA has a very severe peroxide restriction because peroxides lead to early radical formation and knocking in internal combustion engines. A thermal treatment [27] is used for their destruction. Alternative catalytic technology is also patented [28]. The dehydration process to isobutylene (see later) is much more forgiving. Only impurities in the iCl~ boiling range and precursors of such impurities need be removed.

Isobutylene and MTBE Production The dehydration of TBA to isobutylene and water is an endothermic process (AH r = 16.04 kcal/g-mol at 25°C) that is favored by high temperature and low pressure. The equilibrium constant is PiC 4 = Pu20

Kp = — -----

(22)

r TBA

Partial pressures (atm) are plotted in Fig. 1. To achieve 99% conversion at 1 atm, a temperature of 271°C (520°F) must be reached. The materials are in the vapor phase under these conditions. Many inexpensive acid catalysts are available to promote this reaction [29]. Commercially, one has the al­ ternatives of placing catalyst-containing tubes in a furnace, circulating hot gases over tube banks, or using adiabatic beds with reheat furnaces.

99

Propylene Oxide and Glycol

Children Face Tough Issues FIG. 1

Effect of temperature on equilibrium constant of reaction TBA = z'C4 = +H:0 .

The effluent contains unconverted TBA, impurities in the TBA, and water. Specification iCl~ (Table 7) can be produced easily from this mix. Alternative dehydration technology is provided by Huels [30]. The re­ action is carried out in the liquid phase at C H 3 -C H (O H )C H 2Cl + HC1

(32)

(propylene chlorohydrin)

C H 3 -C H (O H )C H 2Cl + 2N aO H or C a(O H ) 2 -----> PO + 2H 20 + 2NaCl

or

CaCl 2

(33)

These units fall into three categories: 1. 2. 3.

Large-capacity units closely integrated with energy-efficient Cl2 plants (e.g., Dow). Smaller, completely depreciated units converted from E O service [50], usually with captive dem and. Units in developing countries operating behind a tariff or transportation barrier, where the local economy does not provide sufficient dem and to justify an oxidation-type plant.

105

Propylene Oxide and Glycol

Types 2 and 3 are gradually being shut down as the plants wear out, w ater effluent laws stiffen, or m arkets change. Occasionally, a new small chlo­ rohydrin plant is built in a rem ote developing country. Dow is said to elim ­ inate CaCl 2 effluent by using N aO H and recycling the resulting NaCl. This technology is described in a num ber of references [51, 52].

N ew Routes to Propylene Oxide and Propylene Glycol The oxidation routes were dem onstrated to be economic in the early 1970s. How ever, since 1972 there have been over 500 references cited in Chemical A bstracts on the m anufacture of PO and PG. Those where a m ajor effort has been made are surveyed here.

Hydroperoxide Processes Texaco has been active in developing a P O /T B A process. They have p at­ ented several molybdenum- [53]* and boron- [54]* containing epoxidation catalysts. There are about 30 references to m olybdenum catalysts [55]. Tincontaining catalysts have been patented by R hone-Poulenc [56], rhenium by Sun [57], and various m etal phthalocyanines by Mitsui Toatsu [58]. Texaco has patented process concepts [28]. O thers have claimed Shell-type (Ti) catalysts [59].

H20 2 Processes Since 1965, there have been over 40 references to use of H 2 0 2 as an epox­ idation agent. The source of the oxidant is ultim ately oxygen, with w ater as coproduct, which does not help pay for the capital and operating costs of making the H 2 0 2. In addition, a substrate must be hydrogenated [60], making the overall stoichiometry the same as a no-coproduct hydroperoxide process: H2 + 0

2

+ C32- -----►PO + H 20

(34)

A catalyst is required to achieve oxidation, and most of the references claim such catalysts. Several acids are cited: formic [61], acetic [62], p ro ­ pionic [63], and sulfuric [64]. Metals used are Ag [65], B [6 6 ], Cu [67], Hg [6 8 ], Mo [69], Pb [70], Pd [71], Pt [72], Sb [73], Se [74], Si in the vapor phase [75], Sn [76], Ti [77], and W [78]. Shell has used fluorinated ketones

[79]. Most H 2 0 2 is currently made by oxidation of a substituted hydroquinone, such as 2-tert-buty\ anthraquinone [60]. An interesting variation by BASF *In this new routes sectio n , only a few o f the pertinent references in each category are supplied because o f their great number.

Propylene Oxide and Glycol

106

[80] uses the raw oxidate for epoxidation (over T i0 2 and S i0 2). This could reduce capital and operating costs. In 1980 Cetus revealed an enzyme-based process for PO [81]. The steps are as follows: o

i

^

enzyme A

Substrate + 0 2 ---------- > H2 0 2 + coproduct additional steps

H2 0

2

(35) marketable > coproduct

+ C 3 - + haloacid enzyme > propylene halohydrin enzyme C

Propylene h alo h y d rin -------- > PO + haloacid

(36) (37)

The following pairs of coproducts were proposed: Substrate

Marketable Coproduct

Methanol Glucose Glucose

Formaldehyde Fructose Furfural

Cetus also stated that the economics of the glucose-to-fructose step in re­ action (35) were attractive enough that it could stand on its own. N either fructose alone nor PO-fructose have been commercialized. W hen secondary alcohols are oxidized, a “ ketone peroxide” is produced. These may be written as R 2 C (O H )O O H or R 2 C 0 H 2 0 2, so they are m ore properly “ alcohol hydroperoxides” or “ ketone-hydrogen peroxide adducts.” BASF claims a route [81] that combines oxidation of cyclohexanol to “ cyclohexanone peroxide,” with subsequent epoxidation using the Shell T iO r S i0 2 catalyst [8 , 9]. The coproduct cyclohexanone must still be used in nylon. How ever, if the cyclohexanol has been derived from the cyclohexyl hydro­ peroxide + propylene epoxidation [49], half as much coproduct is produced.

Peracid Routes Peracids are m ade by oxidation of aldehydes at low tem peratures (0-50°C) and react with olefins to m ake epoxides selectively at m oderate tem peratures (50-100°C) [1]. References are available for PO form ation from peracetic, perpropionic, perisobutyric, perbenzoic, and pertoluic acids. H ow ever, be­ cause of by-product m arkets, only peracetic and pertoluic can be considered [82]. Use of acetaldehyde to m ake acetic acid has been supplanted by m eth­ anol carbonylation, and production of peracetic acid is highly hazardous. Thus, it is not surprising this long-known and w orkable route has not been exploited. The pertoluic routes are m ore interesting. H ow ever, oxidizing the second methyl group to produce terephthalic acid (TPA) is difficult. Exploitation of this route is certainly limited to those who have fully developed TPA

Propylene Oxide and Glycol

107

technology. Separation of the volatile PO from the product acid with m in­ imum loss is a problem with all peracid routes.

Liquid-Phase Cooxidations Liquid-phase cooxidations to propylene oxide and a coproduct fall into three categories: 1. 2.

3.

A n aldehyde plus propylene to produce a peracid in situ A hydrocarbon plus propylene plus variable valent metals to produce hydroperoxide in situ A n alcohol plus propylene to produce H 2 0 2 in situ

All operate under conditions in which allyl oxidation of propylene [reactions (l)-(4 )] is significant. This means that yields of PO from propylene are poorer than two-step processes, and product separations are difficult. Every aldehyde-acid pair tried for the two-step process has also been tried in a one-step process [83]. Hydrocarbons used include ethyl benzene, isobutane, cum ene, and cyclohexane [84]. The only example of an alcohol cooxidation, isopropanol, reported an explosion during the workup [85]!

Coproduction of Acrylic or Methacrylic Acid Three interesting routes in which acrylates are coproducts have been re­ ported. Sumitomo claims cooxidation of acrolein and propylene with a tran ­ sition m etal catalyst [8 6 ]. Halcon claims cooxidation of p-m ethoxyisobutyraldehyde and propylene [87]. The acrylic route must com pete with the facile two-stage vapor-phase oxidation. The m ethacrylic routes obviously require m ultistep systems to get to the oxidant and to crack the resulting ether or alcohol. In all the routes in which isobutyric acid is a coproduct, dehydrogenation to m ethacrylic acid is conceivable (as well as propionic acid to acrylic acid). Asahi [8 8 ] claims cooxidation, esterification of the product isobutyric acid, and dehydrogenation.

Liquid-Phase Oxidation Without Coproduct There are over 40 references to oxidation of propylene according to reactions (l)-(4 ) dissolved in a variety of solvents using a variety of catalysts, som e­ times with cooxidants (e.g., N 2 0 4). However, they all show relatively poor selectivity and produce corrosive effluents that are difficult to work up. Of more interest are the relatively few patents in which propylene is directly oxidized by the catalyst: M z+2+ + C32- -----> M2+ + PO

(38)

108

Propylene Oxide and Glycol

and the catalyst is reoxidized by m olecular oxygen, either in situ or after removal of the propylene and PO: 2M Z+ + 0

2

-----> 2M2+2+

(39)

The actual mechanism may involve direct oxygen transfer: L mM O + C32- -----> L mM + PO

(40)

or w ater oxidation with rem ote reoxidation: L mM z+2+ + H 20 + C32- -----> L mM z+ + PO + 2 H +

(41)

The most extensive work has been done with thallium. Early work by H e r­ cules [89] and R hone-Poulenc [90] produced glycol and glycol acetates from thallic acetate; Teijin [91] produced halohydrins from thallic halides. Asahi produced PO directly in thallic acetate [92] or over supported thallic oxide [93]. A R C O [94] regenerated acetate solutions in alkali and then acidified with C 0 2. Halcon [95] carried out the epoxidation in the presence of C 0 2. Even the m ore practical of these routes require considerable equipm ent to reoxidize the thallium , and none allow sim ultaneous epoxidation and reoxidation. Several patents claim catalytic action under both propylene and oxygen. The catalysts are Ag or Ru [96]. However, the tem peratures are high enough that the radical mechanism [reactions (l)-(4 )] must be suspected. Several papers cite the use of m etal porphyrins at near room tem perature to epoxidize higher olefins [97]. If propylene were to work in these systems, many orders of m agnitude of rate increase would be required for com m er­ cialization. Several references cite direct oxidation with m etal salts, but no indus­ trially practical regeneration is suggested. The metals are Hg [98], Ru [99], Pd [100], A u [101], and Cu [102]. An interesting redox couple is periodate-iodate, which was shown by Dow [103] to oxidize propylene. Reoxidation requires high temperatures if 0 2 is used.

Ring-Closing Routes Several groups have found apparently attractive liquid-phase oxidation routes to propylene glycol (PG ), propylene glycol esters, or propylene car­ bonate. Since PO use in non-PG m arkets m akes up 70% of the PO + PG m arket, these researchers looked for ways to produce PO from derivatives. First, the oxidation routes and then the ring-closing routes are reviewed. R outes to PG and PG esters using thallium were presented earlier [89, 90]. The Halcon tellurium brom ide acetic acid route, which was abandoned for ethylene glycol production [104], will also m ake PG and PG esters [105]. M ontedison claims propylene carbonate production in M n 0 2-iodide systems

109

Propylene Oxide and Glycol

pressurized with C 0 2 and 0 2 [106]; it is not clear w hether M n 0 2 is an oxidant or catalyst. A R C O patented the F eI 2 /C u S 0 4 system, which m akes propylene carbonate [107]. Asahi patented a PdCl2 H N 0 3-acetic acid system for pro­ duction of glycol acetates [108]. It should be noted that all these systems require exotic materials of construction and require the recycle of consid­ erable quantities of catalyst. Propylene carbonate seems to be the easiest to convert to PO. Texaco [109] and others [110] have catalytic technology that works at low tem per­ ature (e.g., 160-240°C). PG and PG acetates are m ore difficult to close, requiring catalysis at 270-400°C. A R C O [111] has technology for PG con­ version. Chem Systems [112] and others [113] have patented PG acetate technology. A nonhalide, nonaqueous process to PC could be commercially attractive.

Vapor-Phase Oxidation All ethylene oxide is m ade commercially over supported silver catalysts in the vapor phase [104]. It is m ore difficult to produce PO in the same way because of the allyl oxidation problem . A recent example [114] reports 0.78% conversion of propylene at 21% selectivity to PO. W ork has also been done with Cu and Fe com pounds. The best results are 35% selectivity at 8 .6 % conversion over a C a/T l catalyst [115].

Halide and Electrochemical Routes The basic problem with the chlorohydrin route is the cost of making chlorine electrochemically. Much of the recent work in the area has been on making PO in electrolytic cells containing halide and through which propylene is bubbled [116]. This work does not answer the basic problem that the oxidant is still being produced by using electrical power. Use of iodine to form propylene iodohydrin could be attractive because iodides can be oxidized to iodine with H 2 S 0 4 or air. Shell [117] patented such a process in 1969. However, the authors have found no work since that date. The rate of iodohydrin form ation is slow because of the low ionization of iodine. However, this is a no-coproduct process, and it is surprising that m ore effort has not been m ade. Lummus developed a process in which TB A was converted to a hy­ pochlorite and used to oxidize propylene [118]: TB A + Cl2 -----> (C H 3 ) 3 C0C1 + HC1 C H 3COCl + H 20 + C32 - -----» propylene chlorohydrin + TB A

(42) (43)

It is hard to see how this is a significant im provem ent on basic chlorohydrin technology.

110

Propylene Oxide and Glycol

Biologic Routes C etus’ work using extracted enzymes in a multistage process was just de­ scribed in H 2 0 2 Processes. From 1980, efforts in the biosynthesis have been shifted to using whole-cell microorganisms. The popular ones are m ethane or m ethanol consuming and produce single-cell protein (SCP) as a by-prod­ uct. A m ong the active research groups are Exxon [119], the University of Pittsburgh [120], Nippon Mining [121], the Bio R esearch C enter (Japan) [122], and the A gricultural University of W ageningen, N etherlands [123]. All these groups are active in efforts to immobilize the microbes. This is by far the most active current PO research activity. Much work needs to be done before competitive economics are achieved.

Properties of Propylene Oxide Physical and Chemical Properties Propylene oxide is a colorless, low-boiling flam m able liquid that is m od­ erately toxic. It has limited solubility in w ater but is miscible with a large num ber of organic solvents. The strained epoxide ring can be opened with a variety of com pounds. In some cases these reactions can occur w ithout a catalyst. M ore often, these reactions are catalyzed by bases and, in some instances, acids. Some physical and therm ochem ical properties are sum m arized in Table 9. A graph of propylene oxide vapor pressure versus tem perature is given as Fig. 2. M ore extensive data and tables are available in the literature. The struc­ ture of propylene oxide, also known as 1 ,2 -epoxypropane or m ethyl oxirane, is commonly represented as H2C------- CH - CH 3 O

Specifications of Product C urrent general sales specifications from a m ajor U.S. m anufacturer are shown in Table 10.

Chemical Properties Propylene oxide is a reactive com pound and is used as a chemical inter­ m ediate for the m anufacture of a variety of end products. M ost reactions involve opening of the epoxy ring: H3C HC

H2C

H3C

\ /

O ( + XY

HC — OX H2 C — Y

(44)

Propylene Oxide and Glycol

TABLE 9

111

Physical and Thermochemical Properties of Propylene Oxide Sourcea

Liquid Molecular weight Boiling point 760 mm Hg, °C 300 mm Hg, °C 100 mm Hg, °C ABP/AP at 740-760 mm Hg, °C/mm Vapor pressure at 20°C, mm Hg Density at 20°C Specific gravity at 20/20°C A(specific gravity)/At at 20-30°C Water solubility at 20°C Freezing point, °C Viscosity at 10°C, cP Surface tension 20°C, dyn/cm 0°C, dyn/cm Refractive index, n2{) Specific heat 20°C, cal/g/°C 0°C, cal/g/°C Heat of vaporization at 34.23°C, cal/g Heat of fusion, kcal/g-mol Flash point, tag-closed cup, °C Vapor Specific heat at 25°C, cal/g-°C Critical temperature, °C Critical pressure, kg/cm2 abs Autoignition temperature in air at 1 atm, °C Explosive limits in air at 1 atm Upper, % vol Lower, % vol. Heat of formation at 25°C, kcal/g-mol Thermal conductivity at 25°C, g-cal/h-cm2-°C/cm

58.08

1

34.23 10.8 -12.16 (minus) 0.038 440.85 0.8299 0.8313 0.00129 39.5 -112.22 0.36

1 1 1 1 1,2 1

25 26.3 1.3665

1 1 1

0.48 0.47 113 1.5614

Ui oc G|

\

p.

\

\ \

_____________ ^

x

— .-C_

a

X

—~—

dAP 0? ' dQ

ORIFICE AP R • V2

i

\/ / X\ \ H 01

X

D NL \ s r03

- j

_ FLOW. Q

i—i—r I AP I

A. SIMULATED COMPRESSOR SOURCE, SHOWING NEGATIVE DAMPING FROM A TO B

02 NEGATIVE SOURCE DAMPING GC

3 V) C /) UJ GC

0.

1 HIGH PIPE D A M P IN G fj

LO W PIPE D A M P IN G

PROGRAMMABLE ORIFICE

DAMPING RESISTANCE

x j-

B. PIPE ACOUSTIC DAMPING IN COMPARISON TO SOURCE NEGATIVE DAMPING F IG . 8

Flow stab ility c rite ria fo r a passive piping sy stem using a h ypothetical flo w -co n tro llin g o r­ ifice to sim u la te a c o m p re sso r head curve.

the orifice drop; that is, PL = P 0 — A P. Note that betw een points A and B, the required orifice resistance is negative; that is, as flow increases, orifice pressure drop decreases. If we attach piping w ith a dam ping line r s as shown in Fig. 8b \ the system is stable because the positive dam ping afforded by piping is greater than the negative dam ping o f the com pressor. For load line r2, however, the system is unstable at the operating point O because negative dam ping R Q2 is greater than load dam ping r2. T his approach was confirm ed using electronic sim ulation, and the electronic system indeed becam e unstable. In addition, the sim ulation frequencies were very close to observed surge frequencies o f the laboratory com pressor for a broad range o f piping dim ensions. Surge frequencies could be accurately predicted using the electronic m odel. The im plication o f course is that surge is a system insta­ bility, dependent only upon head curve and piping configuration. In either the electronic or fluid system , conditions are not stable in the negative dam ping re­

Pulsation in Centrifugal Compressors

201

gion unless external dam ping is sufficient to make total dam ping positive. The system may be m ade up of a centrifugal com pressor, electrical generator, or even a reciprocating com pressor. The instability or surge does not occur because ex­ ternal driving forces (vortex form ation, flow sep aratio n s, and others) are greater than dam ping forces; it is sim ply a result o f inherent instability (actually, bista­ bility), and it will be unstable even when no external (or internal) exciting forces are present. Stall, flow separation, or any o f various possible vortex generators may play an im portant p art in shaping the head curve in the region o f surge, but the sim ulation tests show that surge itself is a rath er different m echanism whose onset is determ ined by system acoustic dam ping and w hose frequency is d eter­ m ined largely by attached piping.

R eactive Piping System s The piping im pedances provided by real piping system s are substantially more com plex than those o f the sem i-infinite, constant diam eter pipe described earlier. Any change in flow diam eter affords reflection points for pulsations, and pipe lengths strongly influence piping resonant frequencies. The net effect is an ex­ trem e variation in pipe im pedance as a function o f frequency, and preferred fre­ quencies or resonances are quite pronounced. C om plex piping system s made up o f various discrete lengths, diam eters, branches, volum es, and constrictions exhibit dynam ic transfer characteristics that am plify pulsations at som e frequencies and attenuate them at others. In this ac­ tion, the piping characteristics are alm ost directly analogous to electrical circuitry or delay lines, in w hich the inertial (inductance) and stiffness (capacitance) ch ar­ acteristics are quite linear but the resistance is nonlinear with flow. In flow ing piping system s, the acoustic resistance is a direct function o f the p V 2 pipe fric­ tional losses. These far overshadow conventional acoustic dissipation effects, such as m olecular relaxations, for the range o f flow s norm ally encountered in industrial piping system s. This distributed im pedance netw ork can am plify low-level pulsation up to as much as a factor o f 100 (i.e. acoustic, Q up to 100 or so), depending on frequency, pipe size, flow, term ination im pedance, and fluid properties. Since the reactive flow im pedances (distributed acoustic inertia and stiffness) o f a piping system are quite linear, conventional acoustic theory can be used to define resonant frequencies, m ode shapes, and filtering characteristics even for rather high flow velocities. The nonlinear dam ping affects only the am plitude or sharpness o f the resonance peaks. Using elem entary physics, resonances o f indi­ vidual piping segm ents can be very sim ply described from “ organ p ip e” reso­ nance theory. To illustrate the influence of finite length piping, consider the standing wave m odes o f a sim ple clo sed -clo sed pipe section (i.e ., capped at each end). Figure 9 shows both the pressure and velocity standing wave patterns at resonance. By dividing pressure p by velocity v, the im pedance z along the pipe is plotted for its lowest (half-w ave) resonance.

Pulsation in Centrifugal Compressors

202

I______________________________________J CLOSED-END PIPE. HEADER, OR VESSEL

PRESSURE

VELOCITY

IMPEDANCE ( - ^ )

FIG . 9

S ta n d in g w ave plo ts o f a c lo se d -e n d pipe reso n an ce show ing v a ria tio n s in p re ssu re , flow , and im p ed an ce as a fu n ctio n o f d ista n ce along the pipe.

From this im pedance plot, it may be seen that reactive im pedance is zero at the point o f pressure m inim um (velocity m axim um ). If a centrifugal com pressor feeds into the center o f such a pipe or vessel, a very low resistive im pedance (or slope) is seen. A lternatively, if the feed-in point is near a capped end, a very high reactive im pedance is seen by the com pressor. From earlier discussions, one m ight expect to illustrate the effects o f reactive im pedances sim ply by changing slope o f the load line. This im plies a situation like that in Fig. 10, w here two loaded lines pass through the operating point O. U nfortunately, the process is m ore com plex because o f phase shift betw een pressure and flow near resonance. At resonance, this shift is 90° (in tim e) for sin­ gle degree o f freedom system s and tends to produce an elliptically shaped orbit rather than a single straight line. If pressure and flow were m easured at a pressure m axim um point in the piping, the load line w ould indeed be nearly a straight vertical line (large p , zero v). At a velocity m axim um , it w ould be nearly hori-

Ui

C HIGH IMPEDANCE ) FEED

CC o

CO CO UJ

POINT

CC CL UJ

" LOW IMPEDANCE FEED POINT

0

CC


, “ A” PUMP DISCHARGE FILTER INSTALLED . FILTER OUTLET

I I / ! /

CL

O

400

2 < z °

I 200

I I I \ I \

•s_ ' I

0 15

20

25

30

35

40

FREQUENCY, Hz FIG . 4

E ffect o f a co u stic filter on p u lsatio n s from a re c ip ro c a tin g trip lex pum p.

Pulsation in Centrifugal Pumps

221

ACC UM ULA TO R

Ol-------------- 1 S U C T IO N — I—

,

PUMP

|----------- ,

]

— i_ , "C" P U M P ]

i-------------- 1

|T--------------------- 1 I

ACCUM ULATOR ,

?

t

r

_ IS

^

^

.

FILTE R fiL r E -

? t-n

HARGINGfcrt^ | 1 ! Z ! 1 !

c h a r g in g G E

lineJ u n e

|

2

S EEAL AL L L IN IN E E ‘‘ S U C T IO N — 1! i

i)

“ ! Jj

e!2 ! S < I 3

z o cr O E N G IN E F U N D A M E N T A L M A X . L O A D C O N D IT IO N M IN . L O A D C O N D IT IO N

I_____________ I__

700

750 E N G IN E S P E E D , rpm

FIG . 12

T orsional v ib ratio n s as a function o f pum p speed at m inim um and m axim um

load

co n d itio n s.

Case 3. Failure of Pump Internals A four-stage centrifugal pum p suffered repeated failures o f the splitter between pum p stages. A detailed field study revealed that the cause o f the problem s was an acoustic resonance o f the long crossover that connected the second-stage d is­ charge w ith the third-stage suction (Fig. 13). The resonant frequency was a half­ wave acoustic resonance.

SUCTION

FIG. 13

A four-stage pump.

Pulsation in Centrifugal Pumps

228

w here: a = speed o f sound, ft/s L = length, ft The speed of sound in w ater is a function o f the tem perature, and at 310°F, the speed of sound w as 4 ,7 7 0 f/s. The length o f the crossover was 5.75 ft. The acous­ tic natural frequency is

/ = 2 0 7 5 ] = 415 HZ

(8)

The acoustic resonant frequency was excited by the blade passage frequency (7 tim es running speed). C oincidence occurs at (415)(60)/7 = 3 ,560 rpm . Pulsations m easured in the crossover show ed pulsation am plitudes o f 100 psi peak to peak. These pulsations were attenuated to the point that at the suction and discharge flanges, the am plitude was less than 10 psi peak to peak. Because the severe vibrations occurred w hen the speed was 3 ,560 rpm for a w ater tem perature o f 310°F, several possible changes could be m ade to elim inate the possibility o f coincidence of resonances during norm al pum p operation. O ne possible change was to reduce the diam eter o f the im pellers and operate the pum p at a higher speed. A nother possible change w ould be to change to six or eight blades to change the blade passage frequency. The im peller diam eter change was the quickest and was carried out in the field, and the failures w ere elim inated.

Conclusion The ability o f centrifugal m achines to am plify o r attenuate piping pulsations has been docum ented in both the field and the laboratory. C oexisting acoustic reso­ nances o f the piping system can further am plify pulsations at frequencies co rre­ sponding to conventional organ pipe resonances o f pipe com ponents or by interacting system acoustic resonances o f m ultiple com ponents. T he net result is that relatively low level excitation sources, such as vortex form ation at jun ctio n s in the piping system s, can be am plified to levels that cause failure of piping and com pressor or pum p internals. Recent research into the m echanism s o f these selfsustaining resonant buildups provides prediction techniques whereby such pulsa­ tion problem s can be m inim ized, and case histories have proven the applicability o f acoustically m odifying pum p and piping internal design to solve system vibra­ tion problem s. M uch o f this article was excerpted by special perm ission from Hydrocarbon Processing, June 1987.

References 1.

W . V o n N im it z , " R e lia b ilit y an d P e r fo r m a n c e A s s u r a n c e in th e D e s ig n o f R e c ip r o ­ c a tin g C o m p r e s s o r I n s t a lla tio n s ,” P r o c e e d in g s o f th e 1 9 7 4 P u rd u e C o m p r e s s o r T e c h ­ n o lo g y C o n fe r e n c e , P u rd u e U n iv e r sity , W e st L a fa y e tte , In d ia n a , J u ly 1 0 - 1 2 , 1 9 7 4 .

Pulsation in Gas M etering Systems

229

2.

C . R . S p a r k s , " P la n t P u ls a tio n s and V ib r a tio n C o n tr o l," p r e se n te d at S o u th w e s t P e ­

3.

C . R . S p a r k s . " A n a lo g S im u la tio n o f C o m p r e s s ib le P ip e F lo w ," A S M E 7 0 -P e t -3 3 .

tr o le u m S h o r t C o u r s e , L u b b o c k . T e x a s , A p r il 1 5 - 1 6 , 19 7 1 .

C E C IL R. S P A R K S J. C . W A C H E L

P ulsatio n in Gas M e te rin g System s

In tro d u ctio n R ecent years have seen an increasing aw areness o f the need for im proved flow m easurem ent accuracy at natural gas m etering installations. In response to these needs, several research program s have been initiated by industry groups, w ith ini­ tial efforts generally concentrating in two areas: (1) developm ent of more accurate coefficient data for the orifice, and (2) developm ent o f im proved m etering con­ cepts to replace the orifice. Both approaches have m erit as long-range objectives, but it is clear that neither approach in itself is sufficient to solve all field m ea­ surem ent problem s. This is in large part because o f the ill-defined and poorly con­ trolled flow conditions and gas ch aracteristics som etim es seen in the field. Until such field conditions are identified and corrections or com pensations are m ade, it is unlikely that any m easurem ent technique can be totally successful in providing the hoped for im provem ent in field m easurem ent accuracy. It is the intent of this article to discuss various nonideal flow conditions that occur at field m etering sites (the so-called system s effects) and to describe the results o f recent research sponsored by the Southern G as A ssociation Pipeline and C om pressor Research Council (SG A -PCR C). This research has been specifically concerned with developing m eans for identifying, defining, and controlling sys­ tem s effects at natural gas m etering installations.

Background Several program s have been conducted in recent years, both in the United States and in the European Econom ic C om m unity, concerned with im proving o rifice co ­ efficient data. Program s sponsored by the G as Research Institute, A m erican Pe­ troleum Institute and A m erican G as A ssociation at the N ational Bureau of Standards L aboratories in G aithersburg, M D , and Boulder, C O , and at the C ol­ orado E ngineering Experim ent Station are notable exam ples o f this w ork. Parallel test w ork in the United K ingdom at British G as and N ational E ngineering Labs,

230

Pulsation in Gas Metering Systems

at G asunie in the N etherlands, and at G az de France have included round-robin com parison testing o f a few standardized orifice sizes. In all cases, the prim ary objective was to substantially extend the orifice data base, to m inim ize data scat­ ter. and to define the seriousness of "facility b ia s" effects that appear to be in­ herent in the various individual test facilities. These tests have been described in a variety o f publications and are not dis­ cussed here, other than to acknow ledge that results have not as yet been com ­ pletely analyzed and disclosed. It is expected, however, that results will serve to enhance the basic accuracy capability o f the orifice and to increase in d u stry ’s con­ fidence in the orifice as an accurate and viable custody transfer meter. Certainly, all questions will not be answ ered by these tests, and interest has already been expressed in expanding the data to larger orifice sizes and w ider ranges of R eynolds num bers and to include sensitivity analyses relating coefficient data to various m echanical param eters, such as plate thickness, tube roughness, and concentricity. M any questions rem ain concerning these tests, but som e things are already known. It is generally recognized, for exam ple, that accuracies obtained in the laboratory are substantially better than those obtainable at typical field installa­ tions. M erely squeezing out the scatter and u ncertainties in laboratory m easure­ m ents will therefore do little to m itigate som e o f the m ajor sources o f erro r in the field. M uch of the disparity between laboratory and field accuracies is o f course a direct result o f system s effects, that is, the less than ideal flow conditions in the field. O f particu lar significance are pulsations, sw irl, profile distortions (caused by inappropriate upstream conditioning and inadequate station design), variations in gas com position, and entrained liquids. W hereas laboratory conditions are of­ ten closely controlled and specifically m onitored to avoid adverse flow condi­ tions, it should be recognized that the resulting m easurem ent m ethods and coefficient data must ultim ately be used in the field, w here conditions in general are neither closely controlled nor m onitored. It therefore seem s im portant to stress that the answ er to im proved field m ea­ surem ent lies not only in im proving coefficient data or developing new flow m ea­ surem ent concepts, but also in studying system s effects as they relate to existing m eters and other m etering concepts. A significant start has already been m ade by the Southern G as A sso ciatio n ’s Pipeline and C om pressor Research Council (SG A -PD RC) in defining pulsation effects and control techniques. W ork has also been undertaken by several other groups on controlling the effects o f sw irl and flow profile distortions. It is im portant that this w ork be continued and acceler­ ated to bring together past results and to answ er im portant questions:

How are m eter errors produced? How severe and com m on are they? W hat is their econom ic impact? How can field m easurem ents be made to identify adverse system s effects in the field? W hat techniques can be used to control, prevent, or com pensate for such errors?

Pulsation in Gas M etering Systems

231

Test section

,

R efe re n c e sec tio n

■ Pulsator

Acoustic filter

Rlower acoustic with filter!

FIG. 1

Southern G as A ssociation flow facility.

Pulsation Effects Pulsation Generation and Control The obvious answ er to pulsation problem s is sim ply to elim inate or reduce the pulsations. W hether pulsations result from nearby com pressors or hunting regu­ lators or are flow induced, the technology necessary to achieve such control is well established [ l - 3 ] , Pulsation control techniques have proven their effective­ ness in im proving m etering accuracy, but several historical questions rem ain in applying this technology:

W here is pulsation control needed, and how can needs be identified in the field? How much control is necessary, and w hat are acceptable pulsation levels at m eter installations? Is the required pulsation control equipm ent econom ically ju stifiab le?

It w as specifically to answ er these questions that the SGA -PCRC program on pul­ sating flow m easurem ent effects was initiated. The first step in this process was to build a pulsating flow research facility (Fig. I ). This is a closed-loop recircu­ lating system incorporating specialized features for generating, controlling, and analyzing pulsating flow conditions typical o f those found at field m eter instal­ lations. The facility layout provides for two m eter runs connected in series. Each run is acoustically isolated from the other and from the com pressor system . W hen pulsations are introduced into one o f the runs, pulsation errors can be directly quantified by com paring inferred flow in the pulsation “ test ru n ” to that in the steady flow or “ referenced ru n .” This approach m inim izes the need for absolute calibration and has been extrem ely effective for identifying system s effect errors.

Pulsation in Gas Metering Systems

232

8 7

6 5

I

4

0.

3

3(0 I

1 o> -3 l 0.1

I III

i» !«! ' «' ! M !i .....*!»

0.2 0.4 0.6

2

Predicted error,

F IG . 2

I II

4 6 8 5 .1 I 2

I

20

I

i I 1 1 II

40 60 80

n= 1

Field lest o f early square-root theory (I 9 6 2 ). (From Ref. 4 .)

Flange Tap Errors A m ajor concern o f the initial testing on this facility was to evaluate the adequacy of existing theories for describing pulsation erro r in the orifice and to evaluate potential errors in its attached secondary system s. A dequacy o f the then existing square-root theory was in serious question because o f previous test results, such as that in Fig. 2 [4]. In this plot from a 1961 study, observed pulsation erro r at an orifice is com pared to that predicted by velocity m odulation ratios as described by square-root error. The poor co rrelation suggests that either o f the follow ing is true:

1.

2.

Square-root theory is incorrect. O ther error m echanism s exist that are o f at least com parable m agnitude.

The first step in the SG A -PCRC program was therefore an experim ental re­ exam ination o f square-root error. Very b riefly stated, the results show ed clearly that square-root theory was incorrect; that is, erro r cannot be inferred from the velocity m odulation ratio

SRE„

I «« 2

T his inequality results from a basic but erroneous assum ption in the theory that the discharge coefficients for each o f the pulsation com ponents un are equal one to another and are num erically equal to the steady flow coefficient o f the orifice. E xperim ents using hot w ires, a fast response transducer, and very short gage lines show ed this assum ption to be entirely false; individual coefficients for the u„ com ­ ponents can in fact vary by as much as 20; 1, depending upon pulsation frequency and the acoustic response characteristics o f the m eter tubes and piping system .

233

Pulsation in Gas M e te rin g S ystem s

.2 c

H

o

'*5 _e

_ o a? o

LU

a:

w «3

oj 5

_

£ >%

d
P

> C/5 O u
e

.2 o

/5

3

c

O 3 3

T03

/5

a

C

1 d 0030

/5

C

JU M a K 01 (£>

d

[I

314

315

Centrifugal Pumps TABLE 6.1

Multistage Pump Configurations3

Shaft Orientation:

Casing Construction: Impellers

Thrust Balance

Tandem Tandem Tandem Opposed

N one Individual Balancing device Opposed Impellers

Horizontal

Segment

—

U U A

Axial Split

Vertical

Barrel

Segment

Can

u

—

—

___

—

—

—

A

U

U

A

U -

A —

—

aU = u sual, A = a va ila b le .

For chemical process service, the usual casing configurations are axially split, barrel, and can (in modern parlance, both barrel and can are known as “double casing”). Segmental casings are rarely used because o f the multiple leakage points inherent in their construction, see Fig. 6.43. Axially split casings are the first choice for horizontal pumps. They are limited to working pressures o f approximately 2000 lb/in.2gauge (140 kg/cm2), temperatures to 600°F (315°C), and liquids whose SG = 0.7 or greater. The last two limitations are recommended by API-610 and may be exceeded if there is prior satisfactory experience with the casing in question. Designs for working pressures greater than 2000 lb/in.2gauge are feasible but economically dubious. Barrel casings make for a more complicated pump but are necessary for service conditions beyond the capability o f axially split casings. Most axially split multistage pumps employ opposed impeller rotors. Figure 6.44 shows a typical pump. The claimed virtue o f opposed impellers is inherent zero thrust with no hydraulic balancing device to wear. The usual opposed impeller arrangement, as shown in Fig. 6.44, does have some

F IG . 6 . 4 3 .

M u lt ip le p o te n tia l leakage p o in ts in a seg m en tal, ra d ia lly s p lit p u m p casing.

316

Centrifugal Pumps

A x ia lly s p lit m u ltis ta g e p u m p w ith opposed im p e lle rs fo r h y d ra u lic a x ia l th ru s t b a lance. (C o u rte s y W o rth in g to n P u m p , D resser In d u s trie s, In c .)

■ELA STO M ER S E A L ON R A D IA L J O IN T

S E C TIO N 'A - A ' J O IN T IN T E G R IT Y F U N C T IO N OF

a

PH - P|_ jk F h

.

fT

'

r r

n

r

\ ~ |

FIG. 6 .4 5 .

P|_ - Pa

I

"71 G A S K E T ON A X IA L J O IN T

The “3-cornered jo in t” ; inherent weakness in axially split casings.

317

Centrifugal Pumps

residual thrust (the product o f internal leakage across the center and end bushings) and its magnitude increases as the clearances increase. Despite this, pumps with correctly sized thrust bearings are entirely serviceable. In units of many stages, say six or more, operation o f the relatively slender rotors is doubtless enhanced by the Lomakin effect at the center bushing. The Loma­ kin effect is hydraulic stiffness produced by differential pressure across an annular clearance. One o f the major difficulties in axially split pumps is sealing the inevitable “3-cornered joints,” see Fig. 6.45. As the pressure drop across such joints increases, so does the difficulty o f sealing them. For this reason most barrel or double casing pumps, whose sole justification is better pressure containment capability, employ radially split elements or inner casings. Opposed impeller rotors are not easily incorporated in radially split elements (the necessary crossovers pose a major problem), thus most barrel pumps employ tandem impeller rotors. Figure 6.46 shows such a pump. In the interests o f simplicity and keeping the bearing span short, all but the smallest tandem impeller rotors employ a balancing device to counteract impeller axial thrust. Three types are used: disc, drum, and stepped drum or combined disc/drum. The classical balancing disc, Fig. 6.47(a), is entirely self-compensating, automatically varying its axial gap to develop a balancing force equal to impeller thrust. To be effective, the disc’s axial gap has to be very small. While this lowers the balancing leak-off flow, thus adding 1 or 2 points to the pump’s

FIG . 6 .4 6 .

D o u b le casing (b a r re l) m u ltis ta g e p u m p w ith ra d ia lly s p lit in n e r e le m e n t. ( C o u r ­ tesy W o rth in g to n P u m p , D resser In d u s trie s, In c .)

Centrifugal Pumps

R1R

UJ

o

J Z

UJ

«

O z

< - j £< < — UJ

S uj o

X -J

»1-0 UJ 13

tz *n

3

tz

o: o o

q UJ ■ a. o_

o o UJ z

2 v~

. :

UJ

o

■ UJ

S < < UJ

ZD

co

q

UJ

r | r^-N PSH A = R ^N PSH A

NPSHR

/ Ql 0 2, Q2' F IG . 7 . 8 .

°

R o ta r y p u m p in te ra c tio n w ith system .

7 .6 . Interaction w ith System

The principle governing rotary pump interaction with the system is the same as that for centrifugal pumps: the operating capacity is given by the intersec­ tion o f the pump and system head (or pressure) characteristics. Figure 7.8 shows such an interaction. What is distinctly different in practice, however, is that the very steep pressure characteristic o f the rotary pump means there is very little change in capacity as the system resistance varies. Again, as is the case for centrifugal pumps, the actual operating capacity can be influenced by the NPSHA. If the NPSHA is equal to or below that required, the actual pump capacity will be lower than given by the rating curve. Figure 7.8 also illustrates this.

7 .7 . Parallel and Series Operation

Parallel operation is quite straightforward; capacities are added at equal pressure differentials to produce a combined characteristic, and the pumps operate at the intersection of that and the system characteristic. The only caution is that the pumps be sufficiently similar to flow share. Figure 7.9

358

Rotary Pumps AP

\ r

(3 )

V

\

VS. \

V A L U E -R IS K

\ \

____ O V E R H E A T I N G

_________\

\

\

\

PUMP A

S.

Qa

H I G H A P , QA

F A L L S TO V E R Y LOW

____ ■ — 1---------

OF

8

DAMAGE

PUMPS A 6 B

\Iu M P B

Qb F IG . 7 . 9 .

0 R o ta r y p u m p s in p a ra lle l.

shows an extreme case where one pump has a “softer” pressure characteristic (lower flow regulation) than the other, thus posing the risk o f low flow through the “softer” pump if the pair is run at too high a pressure differential. Given the basic character o f displacement pumps, their operation in series is not a general practice. Circumstances which may warrant it are driver or pump power limitations and, in those pumps capable o f handling abrasive solids, a need to limit differential pressure for longer wear life. The difficulties with series pumping are two. First, the pumps pass the same flow, and unless their pressure characteristics are fairly soft or their control very precise, small differences in pump geometry or speed can mean wide variations in differen­ tial pressure, see Fig. 7.10. Second, succeeding pumps have to withstand higher suction and working pressures.

7 .8 . Varying Flow

Because their flow regulation is very high, sensible flow variation in rotary pumps can only be achieved by bypassing or varying speed. Bypassing, see Fig. 7.11, involves returning part o f the delivered capacity to the pump’s suction vessel (continually returning flow to the pump’s suction should be avoided since it can cause overheating). The bypass valve has to be sized for the maximum flow likely to be bypassed, and its pressure drop is essentially the differential across the pump. As process flow varies, pump gross capacity is practically constant and pump power changes only by the variation in system resistance. Bypassing, therefore, is not efficient. For small pumps or small bypass capacities, simplicity and ease o f operation offset the energy loss. For larger pumps the energy loss may warrant variable speed operation.

AP

\ PUMPS A a B apa + b

-------------

\ \ PUMP B^\\ I

1

PUMPS A a b' OSTENSIBLY IDENTICAL

\Y PUMP A V

_ Q

F IG . 7 . 1 0 .

FIG. 7.10.

R o ta r y p u m p s in series. series.

R otarv Dumps in

Q b = BYPASS

R E L IE F VALVEA

v

»-

|

Q = GROSS

BYPASS

t1

IJ

O v J

Q

- PROCESS

\ \

PROCESS

V

! /

FLOW

SYSTEM

\

"flo^T*"

C A P A C IT Y

CO NTROL

M

"I

1----------

FLO W

»—J -J T 7O 0 PR RO OCCEtS SS

PUMP FLO W

RANGEl

C A P A C IT Y

y \

\

'

'PUMP

PUMP

*

u

R A T IN G

T O BE BYPASSED F IG . 7 . 1 1 .

F lo w v a r ia tio n b y bypassing.

359

Rotary Pumps

360 AP n

NRATE D

n2

H

N,

/

PROCESS

SLIP T Y P E -i DRIVE LOSSES

r

f

* ----------------- ► FLOW RANGE I J

. . ♦ * I l l VFD LOSSES

r

| PUMP POWER

!_________!

► I PUMP

Q

R A T IN G F IG . 7 . 1 2 .

F lo w v a r ia tio n b y v a ry in g speed.

Varying the pump speed is usually realized by interposing a variable speed drive between the pump and its motor or varying the power supply frequency to the electric motor. Variable frequency drive (VFD), as it’s known, offers greater energy savings and is gaining acceptance now that its reliability has been demonstrated. Variable speed operation is shown in Fig. 7.12; pump speed is simply adjusted to produce the required process flow. Pump power decreases with flow and system pressure. Drive evaluation, however, has to take account o f drive losses. When the process flow range is high, hence the speed ratio is high, the higher losses in a slip-type drive tend to offset the higher capital cost o f VFD. Other means of variable speed drive such as steam turbine, hydraulic, or pneumatic motor or engine are, o f course, quite feasible. They are, however, rarely used in comparison to the simple squirrel cage electric motor.

7 .9 . Pump Selection

Rotary pump development has created myriad variations o f the basic con­ cept. The selection o f a particular variation depends upon the service condi­ tions. The particular variation, in turn, can have an effect on the system and pump rating. Figure 7.13 shows the basic process for making and checking a selection.

361

Rotary Pumps

ROTARY PUMP ( F R O M F I G . 5 .2 )

TYPF. G R O U P B A SED ON LIQUID

S P E C IFIC TY P E BASED ON P E R F . R EQ U IR EM EN T S

SY S T E M /R A T IN G C O R R E C T IO N S B ASED ON S P E C IF IC ____________T Y P E ____________ F IG . 7 . 1 3 .

R o ta r y p u m p selection process.

7 .1 0 . N ature o f Liquid

All the variations o f a rotary pump can be assigned to one o f three groups, based on the nature o f the pumped liquid. Four characteristics o f the liquid serve to define the type groups: Rheological characteristic Viscosity and lubricity Presence and nature o f solids Presence o f entrained gas The three rotary pump groups and the broad nature o f the liquids for which they are suitable are listed in Table 7.1.

TABLE 7.1

Rotary P um p G roups and the Liquids for W hich They Are Used G roup

1. Internal m echanism support and force transm ission. 2 . External m echanism support and force transm ission.

3. Elastom er elem ent in m echanism .

N ature o f Liquid M edium viscosity, N ew tonian, good lubricity, free o f solids and entrained gas M edium viscosity, N ew tonian, poor lubricity, nonabrasive solids, entrained gas (in some designs) M edium to high viscosity, nonN ew tonian, abrasive solids

Rotary Pumps

362

L IQ U ID T O BE P U M P E D

R H E O L O G IC A L C H A R A C T E R IS T IC

N O N - N E W T O N IA N ..

_________ __ GROUP 3

N E W T O N IA N

I

____________ Y FS

E N T R A IN E D G A S

»“

GROUP 2

NO

NON A B R A S IV E *”

v FQ S O L ID S

-------------

—

-------------------------■ A B R A S IV ’ -

GROUP 3

NO

V IS C O S IT Y , L U B R IC IT Y

GROUP 2

^

--------------------------

LO W • “

GROUP 2

H IG H

^ F IG . 7 . 1 4 .

GROUP 1 R o ta r y p u m p ty p e g ro u p selection.

Figure 7.14 illustrates how the most appropriate rotary pump group can be chosen from the liquid characteristics.

7 .1 1 . 7.11.1.

System Corrections Rated Flow

With displacement pumps the usual practice is to rate the pump for the required process flow, i.e., not to add any margin. The virtue o f zero flow margin is that it avoids having to provide means to vary the flow. When process flow can vary or the service is likely to cause rapid pump wear, a flow margin is justified. The amount depends upon the likely flow variations in the first case and the desired time between overhauls in the second.

Rotary Pumps

363 TABLE 7.2

Typical Flow Pulsations

P um p Type (see description)

Flow Pulsation, % o f M ean

Any with inlet dam pener Screw (2 and 3) G ear (helical) V ane Lobe: 2 lobe 3 lobe Helical rotor Peristaltic: 2 roller, planar cam 3 roller

0 0 0 5 25 15 3 100 50

7.11.2. Flow Pulsation While it has long been recognized that some designs o f rotary pumps produce pulsating flow, most system design practice makes no allowance for such a flow. This suggests that in general, rotary pump flow pulsations do not manifest themselves as serious problems with NPSHA or pressure pulsations. The form of the flow pulsations and the characteristics o f the liquid may well account for this. For cases where flow pulsations are considered likely to cause trouble or are suspected as the cause o f trouble, Davidson [1.5] provides data for estimating consequent pressure pulsations and acceleration head. As a guide to potential problems, typical flow fluctuations for the usual pump types from Davidson [1.5] are shown in Table 7.2. Installing an inlet dampener effectively eliminates any problem. Dampener selection should be made by the manufacturer, with full knowledge of the form o f pump flow pulsation, and thus must follow selection o f the pump type.

7 .1 2 . Pum p Type

Figure 7.15 shows the types o f rotary pump commonly used for chemical processing. In Fig. 7.15 the pump types are categorized as sealed (having a shaft through the pressure boundary) and sealless (no shaft through the pressure boundary.) Each o f the categories is divided into the type groups discussed above. Group 2 pumps cannot, by definition, be sealless. Typical performance capabilities o f the types in Fig. 7.15 are shown in Table 7.3. These data are, of necessity, only nominal. Individual designs may exceed the nominal, and design evolution can move the nominal, so the table can serve only as a guide.

Rotary Pumps

364 RO TA R Y PU M PS

SEA LED

SEA LLESS

GROUP 1

GROUP 1

S L ID IN G V A N E

M A G N E T IC D R IV E :

IN T E R N A L G E A R

- S L ID IN G V A N E -E X T E R N A L G E A R

EX TERNAL GEAR T H R E E SCREW

GROUP 2

TW O SC R EW

P E R IS T A L T IC (F L E X IB L E

TUBE)

GROUP 2 LOBE EX TERN A L GEAR TW O SCREW GROUP 3 F L E X IB L E V A N E H E L IC A L R O T O R (S IN G L E S C R E W )

FIG. 7.15.

TABLE 7 .3

G roup Sealed: 1

2

3 Sealless: 1

2

R o t a r y p u m p ty p e s f o r c h e m ic a l p ro c e s s in g .

R otary P um ps for Chem ical Processing: Typical Perform ance Capabilities

Type Sliding vane Internal gear E xternal gear Triple screw Tw in screw Lobe E xternal gear Tw in screw Flexible vane Helical rotor M agnetic drive: Sliding vane External gear Peristaltic

Flow (gal/min)

Differential Pressure (lb/in.2)

Viscosity (SSU)

T em perature (°F)

1,000 1,000 2,000 3,000 4,000 400 2,000 10,000 100 1,250

125 250 300 3,000 2,000 450 300 2,500 30 300

500,000 1,000,000 1,000,000 50,000 1,000,000 1,000,000 1,000,000 1,000,000 100,000 1,000,000

450 650 650 200 300 500 650 850 180 200

60 60 300

15 100 220

N.D. 35,000

N.D. N.D. 180

Rotary Pumps

365

In te r n a l (v a n e -in -r o to r ) p u m p . F IG . 7 . 1 6 .

S lid in g van e p u m p . (R e p rin te d w ith p e rm is sio n fr o m R e f. 1.1.)

7 .1 3 . Sliding Vane

The “displacement volumes” are formed between vanes in the rotor, the rotor o.d., and the casing bore, see Fig. 7.16. Displacement o f the volumes is produced by a profiled casing bore. The vanes slide in their rotor slots to follow the casing bore, contact with the casing bore generally being main­ tained by centrifugal force. Critical clearances exist at the vane ends, the rotor ends, and the vane flanks. Wear is most likely at points o f rubbing contact, i.e., the vane tips and flanks. Because there is internal rubbing contact, sliding vane pumps are Group 1 for liquid types. Conventional sliding vane pumps have a single shaft seal. The rotor is either between bearings, with internal bearings, or cantilever, with external antifriction bearings. For corrosive service or where quick dismantling is required, cantilever rotor construction is usual. Wright [7.2] describes such a line of small sliding vane pumps. Sealless sliding vane pumps are available, see Wright [7.2] and Capuder et al. [7.3], but only in small sizes to date. The arrangement used is magnetic drive; see Chapter 6 for details o f the drive.

7 .1 4 . Gear

Two forms o f gear pumps are in general use: internal gear and external gear. The pumping action of both is essentially the same; the “displacement volumes” are formed by the spaces between gear teeth and the casing, and displacement is produced by gear meshing. Figure 7.17 shows an internal gear pump, and Fig. 7.18 shows an external gear pump.

366

Rotary Pumps

(1) Liquid in le t

(2) Passage o f liq u id FIG . 7 .1 7 .

(3) L iquid o u tle t

In te rn a l g e ar p u m p .

Gear teeth, in external gear pumps, are either straight spur or herringbone (gapless double helical). Lower flow pulsation and noise are cited as justifica­ tion for the more expensive herringbone gear teeth. Gear pump “slip,” hence delivered capacity, is dependent on the clear­ ances at the gear ends and gear tips, with the former the more important; see Trushko et al. [7.4], Torque is split between the gears. Internal gear pumps are Group 1 for liquid types. External gear pumps can be either Group 1 or, with external bearings and gear timing, Group 2. Harvest [7.5] cites less shear o f the pumped liquid as the justification for internal gear pumps. Noting that equivalent external gear pumps are less expensive to manufacture, the use o f internal gear pumps would normally be limited to pumping clean, shear-sensitive liquids. Internal gear pumps have one shaft seal. The driving gear (rotor) can be supported by bearings in the pumped liquid or by external bearings. The driven gear (idler) is supported by a bearing in the pumped liquid. The simplest external gear pump, Fig. 7.19, has one shaft seal and bearings in the pumped liquid. For more severe service, those handling Group 2 liquids (nonlubricating), external gear pumps are made with external bear­ ings and timing gears and, of necessity, four shaft seals, see Fig. 7.20.

FIG. 7 .1 8 .

External gear pum p.

367

Rotary Pumps

FIG . 7 .1 9 .

S im p le e x te r n a l g e a r p u m p ; b e a r in g s in t h e p u m p e d liq u id , o n e s h a f t seal. (C o u r te s y W o r th in g to n P u m p , D r e s s e r I n d u s trie s , In c .)

When corrosion is not a problem, gear pumps are furnished with iron casing, end plates and gears, steel shafts, and either antifriction or sleeve bearings. Sleeve bearing materials are usually bronze, iron (in alkaline liquids), or carbon, often with antimony, for low viscosity liquids. Moderately corrosive conditions are met with pumps o f 316 stainless steel. This alloy is noted for its tendency to gall, thus its success in gear pumps is dependent upon proprietary treatments to reduce galling; see Harvest [7.5]. Taylor [7.6] notes the availability o f small gear pumps in alloys higher than 316. Given the simplicity o f the basic external gear pump, Fig. 7.19, there has always been a keen desire to use it in place o f the more complex externally timed version. Progress has been made on this. To retain the simplicity and avoid any stagnant regions, most the effort has been directed toward materi­ als. Trushko et al. [7.4] determined that in pumps so equipped, sleeve bearing wear governed pump life. Carbon on ceramic or ceramic on ceramic (alumi­ num oxide on tungsten carbide or chrome oxide; see Harvest [7.5]) and

FIG . 7 .2 0 .

E x t e r n a l g e a r p u m p fo r liq u id s w ith p o o r lu b ric ity ; e x te r n a lly tim e d , f o u r s h a f t s ea ls, e x te r n a l b e a r in g s . (C o u r te s y W o r th in g to n P u m p , D re s s e r I n d u s tr ie s , In c .)

Rotary Pumps

368

PTFE (in small pumps; see Ref. 7.7) are reported as successful bearing materials. End clearance wear follows bearing wear as the determinant o f pump life. Hardened iron (treatment with nonmetallic carbide, nitride, and sulfide; see Ref. 7.7) is one approach. Nonmetallic “wear plates” is another, nonmetallic gears a third. Gear failure occurs in one o f three ways: end face wear, tooth breakage due inadequate strength, or tooth breakage following gross wear. Nonmetallic gears tend to alleviate the first, but strength can be a problem. Reference 7.8 reports success with carbon fiber reinforced polyphenylene sulfide after breakage problems with glass fiber reinforced fluoropolymer. Koster [7.9] notes that the Hertzian stresses in gear pump teeth can be very high (100,000 to 150,000 lb/in.2 in typical motive power pumps), and that tooth survival is dependent upon elastohydrodynamic lubrication (EHL). Liquid viscosity, its pressure-viscosity characteristic, and the surface finish o f the parts all influence EHL. The significance o f this for chemical processing is a need for caution in application lest rapid tooth wear be a problem. For zero leakage, magnetic drive external gear pumps are available. Such pumps are, o f course, essentially for Group 1liquids, subject to the comments above. As is the case with sliding vane pumps, sizes are presently limited. Nasr [6.33] attributes the limitation to the higher torque required by gear pumps. When the liquid pumped requires heating in the pump, larger size casings can be furnished with an integral jacket. Smaller pumps are available with bolt-on jackets; see Ref. 7.6.

7 .1 5 . S c re w Pumps

Designs employing one, two, and three screws are in use. The single screw pump, more generally known as “helical rotor” or “progressive cavity,” is really a different class o f pumps (Group 3 for liquid type; see Fig. 7.15) and will be discussed separately. In that sense, the pumps to be discussed here are actually “multiple screw.” Following on the fundamental operating principles set out at the begin­ ning o f this chapter, multiple screw pumps operate as follows: The “displacement volume” is opened at the suction as the counter-rotating screws unmesh. Subsequent meshing o f the screws produces a “displacement volume” bounded by the thread flanks, the thread roots, and the pump casing. Continued rotation o f the screws translates the “displacement volume” to the pump discharge. At the pump discharge the volume is displaced by meshing o f the screw ends. Figure 7.21 illustrates the idea.

Rotary Pumps

369

1=0

t= >

(a) F IG . 7 . 2 1 .

(b)

S crew p u m p tran slates flo w a x ia lly ra th e r th a n r a d ia lly as in a n e x te rn a l gear pum p.

Leakage, or “slip,” in screw pumps flows to succeeding “displacement volumes.” Since displacement volume geometry is essentially identical, and neglecting any changes in liquid characteristic through the pump, the pres­ sure rise across the pump (purely a function o f system resistance) will tend to that which balances leakage between the “displacement volumes.” The actual pressure rise distribution at any instant is close to equi-stepped, see Fig. 7.22. This characteristic o f screw pumps leads to the concept o f so much pressure rise per “closure” or “stage.” Brennan et al. in Ref. 1.1 cite typical values o f 125 to 150 lb/in.2 with normal running clearances, and up to 500 lb/in.2 with minimum clearances. The import is that the length o f the pump is somewhat dependent upon the required pressure rise. Zalis [7.10] draws attention to the need for at least two screw turns to avoid “erratic performance and excessive slip.” A by-product of the gradual pressure build up, one coupled with relatively low flow pulsations and low internal liquid velocities, is low noise. Wegener et al. [7.11] report a “medium size” screw pump having a sound level o f 57 dB(A) while running at 2900 r/min and 1450 lb/in.2 gauge discharge pressure.

100 t. Wl 75 50 25

-J-0

IN L E T P R E S S U R E ROTO R L E N G T H (C L O S U R E S )

FIG. 7 .2 2 .

Pressure rise along a screw pum p.

UJ » g5/5 t: 3 O U d E 3

a

o 0

1 "ftj I

CO

K d LL

379

Rotary Pumps TABLE 7 .4

Rotor and Stator Materials for Various Services

Material

Part Rotor

Stator

Nitrided steel Chrome plated 316 Monel or high nickle-molybdenum alloys Natural rubber Nitrile rubber Cast urethane Hypalon3 Vitom

Service Abrasive, noncorrosive Corrosive, with or without abrasives Highly corrosive

General Oil, fats, effluent Abrasive, aqueous slurries Mineral acids, oxidizing chemicals Aliphatic and aromatic hydrocarbons, high temperature

“R eg is tere d tr a d e m a r k o f D u P o n t.

involving a vacuum at the suction, flow is sometimes reversed to be toward the drive so the seal is at discharge pressure. Power is transmitted from the driver via a drive shaft. Two antifriction bearings support the drive shaft, and one o f them also carries the rotor’s hydraulic axial thrust. In some arrangements either bearing can act as the thrust bearing. Casing materials range from iron through 316 stainless steel to higher alloys, depending upon corrosiveness and pressure. For the rotor and resilient stator, Goodchild [7.14] suggests the information in Table 7.4. “Flexible” drive shafts must remain corrosion-free to realize their design endurance strength. Wilson [7.15] notes the use o f “impermeable polymeric” coatings to achieve this. For high viscosity liquids, where gravity flow into the pump is a problem, helical rotor pumps are available in the so-called “open throat” configura­ tion, see Fig. 7.35. In this arrangement the suction opening is oversize and an auger serves to ensure that each o f the pump’s cavities is completely filled. As is the case for multiple screw pumps, helical rotor pumps are not currently available in sealless form. Some work has been done on wet pit versions, but pumps so constructed are not yet in regular service.

7 .1 9 . Peristaltic Pumps

The peristaltic or flexible tube pump is unique in that it’s the only sealless rotary pump with solids handling capability. Maynard [7.17] laments that despite this uniqueness, the device is underutilized.

380

Rotary Pumps

G w ufOf

JJJ

1 [

/

»

t J a r f ^ ==^ ,

26

4A

3018'— * 4 P

301 A-1

N

—

»

307K

In d iv id u a l c y lin d e r b lo c k w ith p r o v is io n fo r v a lv e r e m o v a l th ro u g h m a n ifo ld s. (C o u r te sy W o r th in g to n P u m p , D r esse r In d u stries, In c .)

8.5. Various designs incorporate refinements to improve valve guidance and sealing, see Fig. 8.46. The stuffing box houses the plunger seal, see Chapter 11, and by way o f the seal and throat bushing, serves to guide the plunger. Stuffing boxes are separate; integral designs would complicate cylinder block manufacture. As with piston pumps, see Section 8.19, plunger pumps can be arranged with a surge leg for pumping slurry. The need to do so, however, is less given the plunger pump’s slurry handling capabilities.

TABLE 8 .5

Typical Reciprocating Pump Pressures3

Valve Type

Pressure (lb/in.2gauge)

Plate (disc) Skirt (wing) Ball Elastomer insert “Sou rce: In g erso ll-R a n d ; B u se in [1.1].

5,000 10,000

30,000 2,500

Reciprocating Pumps

420

(a ) o u tsid e g u id e d p la te valve

(b ) o u tside g u id e d s k irt valve

F IG . 8 . 4 6 .

O u ts id e g u id e d p la te a n d s k irt valves. (C o u rte s y W o rth in g to n P u m p , D resser In d u s trie s , In c .)

8 .2 1 . Diaphragm Pumps

The prime virtue o f diaphragm pumps is the absence o f a seal; the pumped liquid is contained and displaced by a flexible diaphragm. A secondary advantage is the ability to run dry. Figure 8.19 shows that there are two distinct classes o f diaphragm pump: direct acting and power. Direct acting pumps have the motive fluid, usually compressed air, applied to the drive side of the diaphragm. The pumps are duplex, see Fig. 8.47.

Reciprocating Pumps

421

With compressed air as the motive fluid, the pressure and power o f direct acting diaphragm pumps is limited. Despite this limitation, their sealless feature means wide usage for handling corrosive, toxic, or viscous liquids. Figure 8.47 shows a typical direct acting diaphragm pump. Its major liquid end components are a pump casing, diaphragms, and suction and discharge valves. Pump casings are made of metals and nonmetals. Metals include iron, aluminum, 316 stainless steel, Alloy 20, and Hastealloy 20. As with centrifu­ gal pumps, nonmetals are replacing exotic alloys for corrosive service. Early pumps were constructed in GRP (graphite reinforced polymer). Nystrom and Larkin [8.3] report the use of polyvinylidene fluoride (PVDF) for pumps handling liquid chlorine and sulfuric acid. The usual construction material for diaphragms is fabric reinforced synthetic rubber. Materials include neoprene, Buna N, butyl, and Viton. Teflon is used either as an overlay on a conventional diaphragm or as the diaphragm itself. Three forms o f valves are used: poppet, ball, and flap, in order o f increasing tolerance o f viscosity and solids size, see Fig. 8.48. Valves are elastomer or elastomer faced. Power diaphragm pumps employ a power pump acting on a hydraulic system to actuate the diaphragm. Figure 8.49 illustrates the principle. These designs are intended to realize the virtues o f power pumps, piston or plunger, while eliminating the seal. This they do, but at higher initial cost and at some expense to volumetric efficiency, the effect o f hydraulic liquid compressibil­ ity. While a diaphragm pump is hermetically sealed, there is the problem of leakage if the continually flexing diaphragm ruptures. The problem is addressed in two steps. First, design has been refined to yield diaphragms whose service life is predictable, thus allowing scheduled replacement prior to rupture. Dailey [8.4] cites the following examples:

M aterial PTFE Metal

Pressure (lb/in.-gauge)

Temperature CF)

Life(h)

5,000 10,000

300 390

20,000 8,000

Vetter and Hering, in Ref. 1.2, note that PTFE diaphragms are tolerant of scratches (notches) but metal diaphragms are not. When incidental leakage cannot be tolerated, a second diaphragm is introduced. Two arrangements are used: double diaphragm and “sandwich” diaphragm. Both have liquid between them, but in the latter the liquid is at atmospheric pressure unless one o f the diaphragms ruptures. Should that happen, a pressure switch senses the pressure rise. The usual power diaphragm pump arrangement is shown in Fig. 8.49; the liquid end is connected directly to the diaphragm housing. For pumping

Reciprocating Pumps

422

A li SUPPl^MCSSUHE

/s i \

F IG . 8 . 4 7 .

i jlb ilfm s

00

I I

1

!i p 3 'i! 1/51^1 I

| | N

> —~


5

p y

O

W

cz

O


u.

& < £

, cl tt

C

=

3

co

1 & »O Z < Z

^

S c r o mi

° ------------------------------

cr 11 a

3

Z

§

& < i Z < 2

< i

.

2

8

_ g £

t 5 O o _ uj

f
LU CL Z X

c H

tP

E

Q

a u ' CO

d C £ 3

S £ 8 £ £ £ u. u- u- ct

z -J


-------------------

P

8

|

S

a.

x

: ______________________________ t

----------------------------------------------------------

_

a-

_____________________ *

^

S

c M U


>

_

------------------------------— —

—

o

s

0

K

l

S y

>

CD

S t

5i

a UJ cr O

?

s e g n

O

„■ I-

i i

I 8 _o_____________________________________________

l

E

_

1 1 ° 2 !

H -

i

a

i

o ®

*

a X

»

3

Pt- aUJ H

°u °

u.

&

o

o p O
C3

O

| < 0 n

~

2

w lu

u

» s 1 ~

O

t

-j < c

M

o I

J

it

“

S

? ° o « " 7 1

l

0 2

§ o

J 3 s « o

.

1

5

I f

2 f S S t < -

t

s

o

n L s z c o

»1