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1995 Optimising alumina feeders in aluminium smelting pots J. P. Kissane University of Wollongong
Recommended Citation Kissane, J. P., Optimising alumina feeders in aluminium smelting pots, Doctor of Philosophy thesis, Department of Mechanical Engineering, University of Wollongong, 1995. http://ro.uow.edu.au/theses/1589
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OPTIMISING ALUMINA FEEDERS IN ALUMINIUM SMELTING POTS
A THESIS SUBMITTED IN PARTIAL FULFILMENT OF THE REQUIREMENTS FOR THE AWARD OF THE DEGREE OF DOCTOR OF PHILOSOPHY
FROM
UNIVERSITY OF WOLLONGONG
BY
J.P.KISSANE, B.E., MTEAUST, CPENG
DEPARTMENT OF MECHANICAL ENGINEERING
1995 This is to certify that this work has not been submitted
for a degree to any other university or institution.
J.P.Kissane ABSTRACT
Crustbreak/feeding systems in prebake aluminium smelting pots are critical to efficient pot control as they contribute in a large way to the alumina balance in the pot. Inefficient feeder and pot operation may cause emission of global warming gases, severe cost penalties and may affect the health and safety of operating personnel This study investigated the causes of failure and the performance of existing designs. Most of the research was conducted at Portland Aluminium (Victoria) and contact was made with about 30 smelters world-wide.
Methods have been developed to improve shot size accuracy, reduce costs, increase crustbreak/feeder life and improved the safety of operators and tradespersons. New designs were developed and two were subsequently patented in Australia, New Zealand South Africa and the U.S.A.
Over 30,000 measurements were taken of dose mass on 7 different feeder designs. Changing the shape of the dosing unit to accommodate better alumina flow rate reduced the standard deviation of the shot size to 0.6% of shot mass for one of the patented designs and 1.9% for the most commonly used pot feeder (compared to an original 7- 14%). These figures compare with 2.5% for the Aluminium Pechiney independent design.
In the investigation into the mechanism of plunger wear and optimising materials, over 10,000 measurements of plunger diameter were taken during feeder maintenance and over 2,500 were taken in 3 monthly measurements in operating pots. Some 19 materials were tested. Plunger life at Portland increased from 1 year to 8 years (for plant scale operation) and trials in operating pots achieved life up to 15 years for one design.
Feeder life at Portland improved by over 80% and repair costs reduced by $1.4m per annum. Feeder life improved from 9 months to 55 months. Many recommendations have been shared at many smelters which has led to significant benefits in improving the existing crustbreaker/feeder units used across the world. ACKNOWLEGDEMENTS 1
ACKNOWLEDGEMENTS
When people consider improving pot feeders, they often think the objectives are to improve life and reduce maintenance costs. These are desirable, but the main benefit is to improve the operation of the pot. A repeatable alumina dose is critical to control of pots, so a good feeder helps the people who control the pot to optimise its operation and reduce generation of potentially global warming gases. Because the feeder is only one part of the equation in pot operation, it is not clear as to how good a feeder needs to be to make a significant effect on how well a pot runs; but poor feeders make pots run poorly.
A long life feeder has a significant effect on the exposure of people to falls, heat, dust and fumes, so some of the main customers in this investigation were the operators and tradespersons. These latter two groups never lost sight of what was the targets were and that ultimately this would help them to work safer in a better environment. These were major incentives to continue when the going got tough.
To complete a study such as this, a team of people need to work together. The bulk of the "hands on" work was done by operators in Portland potrooms and tradespersons from Kempe Project Engineers, Parker Hannifin and Portland Air and Hydraulics. Their assistance and dedication to the strange and repetitive jobs they were asked to do were invaluable in unmasking the vagaries of pot feeders.
Although many people assisted in some way, it would be inappropriate for me not to thank Mike Wakefield for contributing a large part to the data collection on plunger wear. His patience is exemplary. Thanks also to all the "hands on" people: Phil Harvey, Geoff Dover, Doug Lucas, and Kevin (Bronco) Bevis for co-ordinating and recording many tests; Brett Bampfield and Robert Logan for their often hot work on operating pots; Jack Cholewski for co-ordinating the shot size testing and design ACKNOWLEGDEMENTS II changes; Neville Stutchberry for conducting the tests for the pneumatics investigation; Greg Mansfield for the tedious hot job of monitoring the cylinder trials.
The hardest job probably was the tedious brain-deadening operation of the shot size testing...about 30,000 individual mass determinations. Tony Cannon, Sam Kelly and Doug Lucas mainly had this unenviable task. Maybe the most frustrating job was for Val Ellis and Cathy Enscoe for trying to read my writing and translating it into my earlier reports.
There were also a number of people who tried valiantly to improve my engineering knowledge: Alf Jones, Gary Nauer, and Greg Noonan for trying to teach a chemical engineer something about mechanical engineering and pneumatics; Matthew Langmaid and Colm Fitzpatrick for their assistance in turning the A3 feeder ideas into reality; Hans Kempe and Hugh Stark for their valuable advice and knowledge on metallurgy and fatigue; Andrew Morphett for his assistance as a process and statistics advisor; and Fethon Nahoum for his willing help in data analysis and for optimizing feed control so that the hardware improvements were not wasted.
Peter Arnold (my PhD supervisor) was a ready mentor in this development and gave me the inspiration and direction to complete it. To read this document cover to cover is a feat, but to read it several times and to find something to write red comments about on every page was indeed an accomplishment. The quality of the final version is a testament to his patience.
I would also like to thank my opposite numbers from other plants for discussing on phones and faxes the feeder improvement project. Their time and patience during long telephone calls has been appreciated. Probably Telecom and AT&T profits have dropped now the project is complete. The success at other plants is a credit to your teamwork and dedication. Although I helped place a bit of focus on opportunities for improvement, the hard work in the plant is what makes results. ACKNOWLEGDEMENTS m Portland and Alcoa have been most generous in allowing me to use data and facilities to conduct this research. In particular, I want to thank my Potroom bosses...Trevor Minchinton, Gerard Waller and Trevor Adams. They still supported me even at times when results were not coming as fast as they would have liked. The vision of the late David Judd and Geoff Hayward to allow people such as myself to expand our horizons has given me opportunities to learn things and to influence the industry which seldom occurs for engineers. The support of Alcoa USA Primary Metals people Keith Isakson, Bob Seymour and Dick Taylor gave me the tenacity to go on and ensured that others contributed to and benefited from feeder research.
Lastly, I would like to thank my wife Linda and my children Patrick, Anne, Katie and Ben. Ben did a great job in cleaning up the scanned images which took a lot of time. My family listened patiently to all my "feeder stories" and supported me when I needed it, and often lived without me (physically or mentally) on many occasions. It has not been an easy job putting up with me on one of my crusades.
That's a lot of people, but 1 had a lot of help. Thanks.
JimK. NOTE TO THE READER IV
NOTE TO THE READER
Due to the confidential nature of some of the data presented in this document, this thesis is restricted from public access except under written approval from the Technical Manager of Portland Aluminium.
The whole of this work was initiated, controlled and co-ordinated by the author, but considerable input has been given by others. Most of the test work was conducted at Portland Aluminium by the author, operators, and tradespersons. Major contributions that were the work of others has been specified in the text with the name of the person
involved.
To simplify references to plants, any plant specified has been coded throughout the text as follows:
T integrated feeder plant D independent feeder plant
Appendix 10 lists the plant names against the codes.
To assist with brevity and confidentiality, data or opinions of individuals from various plants or organisations have not been cited unless the comment is critical to the argument. However, personal correspondence and drawings are available if required to validate comments. Data from smelters or from organisations have been taken first hand from a suitable representative of that smelter or organisation. If assumptions or interpretations of the work of others are made, they have been noted accordingly.
J.P.K. TABLE OF CONTENTS V
OVERVIEW Page Abstract
Acknowledgements I Note to Reader IV Overview V Table of Contents VI Index of Figures XV Index of Tables XIX Index of Appendices XXII Glossary XXIU
Chapter 1 Introduction and Literature Survey 1 Chapter 2 Shot Size Analysis 61 Chapter 3 Dose Delivery 119 Chapter 4 Pneumatics 143 Chapter 5 Plunger Wear 169 Chapter 6 Plunger Buildup Analysis 230 Chapter 7 Pneumatic Cylinder Investigation 254 Chapter 8 Extending Feeder Unit Life 285 Chapter 9 Cost Reduction 328 Chapter 10 Feeder Development 354 Chapter 11 Impact of Research on Alcoa and Other Smelters 373 Chapter 12 Summary Of Results 391
References 395
Appendices TABLE OF CONTENTS VI
TABLE OF CONTENTS
Page
CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY
1.1 WHY ALUMINA FEEDER TECHNOLOGY is IMPORTANT 1
1.2 LITERATURE SURVEY 7
1.2.1 Aluminium Smelting Technology. 7
1.2.2 Objectives of Alumina Feed Control 13
1.2.3 Initial Prebake Feeding Systems 17
1.2.4 Integrated Feeders 19 1.2.5 Independent Feeders 24
1.2.6 Comparison ofIntegrated and Independent Feeder Designs 27 1.2.7 Conversionfrom Integrated to Independent Feeding 29
1.2.8 Continuous Feeders 33
1.2.9 Aeration Feeders 34 1.2.10 Computer Control. 36
1.2.11 Pot Alumina Supply Systems 38 1.2.12 Plunger Developments 39 1.2.13 Dissolution of Alumina 42
1.2.14 Alumina Dose Design and Accuracy 44 1.2.15 Pneumatic Air Cylinders. 45
1.2.16 Pneumatic Systems 50
1.2.17 Other Feeder Developments 52
1.3 DEVELOPMENT OF THIS RESEARCH 54
1.3.1 Research Objectives 54
1.3.2 Optimizing Existing Feeder Designs 54
1.3.3 New Feeder Designs 56 1.3.4 Across Alcoa Optimisation 58
1.3.5 Document Layout 59
CHAPTER 2 SHOT SIZE ANALYSIS
2.1 SUMMARY 61
2.2 INTRODUCTION 62 TABLE OF CONTENTS VII
PAGE 2.3 PROCEDURE 65
2.3.1 Dwell Time 65 2.3.2 Dosing Unit Shape 67
2.3.3 Stroke Time 69
2.3.4 Validity of Off-Site Testing. 69
2.4 OFF-SITE TESTING OF SHOT SIZE 73
2.4.1 PortlandAEDD 73
2.4.2 T6 Integrated Feeder 78
2.4.3 Tl 1 Integrated Feeder 80
2.4.4 Portland A2 Feeder 80
2.4.5 A3 Feeder and Sequential Feed. 84
2.4.6 Decision on Use ofSpool Inserts 85
2.5 SENSITIVITY ANALYSIS ON SHOT SIZE 87
2.5.1 Bulk Density of Alumina 87
2.5.2 Flow Through Inaccuracies in Dosing Systems 88
2.5.3 Effect of Springs 89
2.5.4 Effects of Cylinder Design on Shotsize 91 2.5.4:1 Description of Cylinder Operation 91 2.5.4:2 Effect of Cushioning on Shot Size 91 2.5.4:3 Effect of Seal Material/Quality 96 2.5.5 Effect of Dwell Time 99
2.5.6 Effect of Height of Alumina in the Hopper 99 2.5.7 Effect of Insert Stroke Length 99
2.5.8 Effect of Spool Manufacture Quality. 100
2.6 ON-SITE TESTING OF SHOT SIZE 102
2.6.1 Til-ColdPots 102
2.6.2 Portland - Cold Pots. 102
2.6.3 Portland-Operating Pots 103
2.6.4 Portland - A3 Feeders 108
2.7 EFFECTS OF PNEUMATICS ON SHOT SIZE 109
2.8 DIRECT FEED INTO THE HOLE Ill 2.9 FACTORS AFFECTING SHOT SIZE ACCURACY IN POTS 115
2.10 MAJOR FINDINGS FROM SHOT SIZE ANALYSIS 117
2.11 RECOMMENDATIONS FROM SHOT SIZE ANALYSIS 118 TABLE OF CONTENTS VIII
Page
CHAPTER 3 DOSE DELIVERY
3.1 SUMMARY 119
3.2 INTRODUCTION 120
3.3 IMPACT OF ANODE COVER ON HOLE BLOCKAGE 121 3.4 CRUST BREAKAGE TECHNIQUES 124
3.4.1 Kinetic Energy 127 3.4.2 Pressure 127
3.4.3 Stroke Time to Break the Crust 128
3.5 BLOCKED FEEDER SURVEY 131
3.5.1 Examination of Blocked Feeder Survey Data 131
3.5.2 New Set Carbon 132 3.5.3 Repeaters 135 3.5.3:1 Dwell Time Tune-out 135 3.5.3:2 Lack of Supply Pressure 136 3.5.3:3 Back Pressure 136 3.5.4 Low bath 138
3.5.5 Pointed Plungers 138 3.5.6 Ore Leaks 140
3.5.7 Feeder Location 140 3.5.8 Minimum Dwell Time Determination 140
3.6 RESULTS OF IMPROVEMENTS TO BLOCKAGE PREVENTION 141
3.7 MAJOR FINDINGS FROM DOSE DELIVERY INVESTIGATION 142 3.8 RECOMMENDATIONS FROM DOSE DELIVERY INVESTIGATION 142
CHAPTER 4 PNEUMATICS
4.1 SUMMARY 143
4.2 INTRODUCTION 144
4.3 GENERAL OBSERVATIONS ON PRESSURE TRACES 147
4.4 METHOD USED TO SLOW FEEDERS 153
4.5 RESULTS 157
4.5.1 Location 157
4.5.2 Feeder Age 157 4.5.3 Mufflers 158 TABLE OF CONTENTS LX
Page
4.5.4 SpeedChange 159 4.5.5 Dual Dwell Time 160 4.5.6 Pressure Required to Break a Hole 161
4.6 AIR SUPPLY OPTIONS 163
4.7 GRADUAL INCREASE IN DWELL TIME 167
4.8 MAJOR FINDINGS FROM PNEUMATICS INVESTIGATION 168
4.9 RECOMMENDATIONS FROM PNEUMATICS INVESTIGATION 168
CHAPTER 5 PLUNGER WEAR
5.1 SUMMARY 169 5.2 INTRODUCTION 170 5.2.1 Pattern of Plunger Wear. 170
5.2.2 Wear Rate Measurement. 171 5.2.3 Theory of Plunger Wear Mechanism 173
5.2.4 Procedure for Analysis of Plunger Wear Mechanism 176
5.3 VARIABILITY OF PLUNGER WEAR RATE 178
5.3.1 Overall Observations 178 5.3.2 Effect of Single Batch Variability on Plunger Wear Rate 179
5.3.3 Effect of Batch-to-Batch Variability on Plunger Wear Rate 183
5.3.4 Effect of Diameter on Plunger Wear Rate 184 5.3.5 Effect ofBetween-Pot and Between-Plant Variability on Plunger Wear Rate 187
5.3.6 Effect of Within-Pot Variability on Plunger Wear Rate 189
5.3.7 Effect of Location Within a Pot on Plunger Wear Rate 191 5.3.8 Recommendations for Data Analysis of Plunger Wear Rates 192
5.4 TESTING CORROSION THEORY 193
5.4.1 Wet Time 194 5.4.1:1 Wet Time per Crustbreak 194 5.4.1:2 Crustbreak Frequency 197 5.4.1:3 Cumulative Wet Time 201 5.4.2 Penetration Distance in Bath 202
5.4.3 Bath Temperature 206
5.4.4 Plunger Preheat 208
5.4.5 Anode Effect Control Strategy 209
5.4.6 Plunger Mass 210
5.4.7 Plunger Buildup 210 TABLE OF CONTENTS X
Page 5.5 TESTING EROSION THEORY 211
5.5.1 Plunger Speed 211 5.5.2 Crust Hardness 213
5.5.3 Plunger Hardness 213
5.5.4 Plunger Shape 215
5.6 PLUNGER MATERIAL 216 5.6.1 General Smelter Experience 216
5.6.2 Kaiser Plunger Wear Experiment 218
5.6.3 Portland Experience 220 5.6.3:1 Cast Iron 220 5.6.3:2 Inconel 222 5.6.3:3 Niresist and Incoloy 222 5.6.3:4 Stainless Steels 223 5.6.3:5 Silicon Carbide 225 5.6.3:6 Alloy Coated Cast Iron 226 5.7 CONCLUSIONS ON PLUNGER WEAR MECHANISM 226
5.8 MAJOR FINDINGS FROM PLUNGER WEAR ANALYSIS 229
5.9 RECOMMENDATIONS FROM PLUNGER WEAR ANALYSIS 229
CHAPTER 6 PLUNGER BUILDUP ANALYSIS
6.1 SUMMARY 230 6.2 INTRODUCTION 231
6.3.1 Dag Frequency 235
6.3.2 Plunger Type 235
6.3.3 Dags Removed. 235
6.3.4 Anode Effects 236
6.3.5 Liquid Levels 236
6.3.6 Location in Pot 236
6.4 CONCLUSION ON CAUSES OF DAGS 239 6.5 FACTORS AFFECTING PLUNGER TEMPERATURES 240
6.5.1 Pot Temperature 240
6.5.2 Plunger Temperature Over Time 240
6.5.3 Hole Size and Flame 242
6.5.4 Number of Holes Open 242
6.5.5 Plunger Length and Penetration 243 TABLE OF CONTENTS XI
Page 6.5.6 Proximity ofPlunger to Hole 244 6.5.7 Wet Time 246
6.6 DAG PREVENTION AND CONTROL 249 6.6.1 Plunger Coolers 249 6.6.2 Liquid Level Control 249 6.6.3 Plunger Penetration Depth 251 6.6.4 Pot Gas Venting 251 6.6.5 Reduced Anode Cover Depth 252
6.7 MAJOR FINDINGS FROM PLUNGER BUILDUP ANALYSIS 253
6.8 RECOMMENDATIONS FROM PLUNGER BUILDUP ANALYSIS 253
CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION
7.1 SUMMARY 254
7.2 INTRODUCTION 255
7.3 CYLINDER TRIALS 254 7.3.1 Introduction 254 7.3.2 Procedure 254 7.3.3 Results. 259 7.3.3:1 Sensitivity Analysis 259 7.3.3:2 Teflon Versus Viton Seals 263 7.3.3:3 Temperature 263 7.3.3:4 Speed 264 7.3.3:5 Eccentricity 264 7.3.3:6 Grease 265 7.3.3:7 Piston rod/shaft connection 265 7.3.3:8 Top cushion length 265 7.3.3:9 Spring design 265 7.3.3:10 Deceleration rate 266 7.3.3:11 Air leak rates 266 7.3.3:12 Insulation 268 7.3.4 Cylinder Trial Summary 268
7.5 LUBRICATION INVESTIGATION 254 7.5.1 Introduction 254 7.5.2 Procedure 257
7.5.3 Kepner Tregoe Decision Analysis Procedure 258 TABLE OF CONTENTS XII
Page 7.5.4 Discussion 260 7.5.4:1 Benchmarks 260 7.5.4:2 No Lubrication 261 7.5.4:3 Lubricant Options 263 7.5.5 Conclusions and Plant Scale Results 264
7.11 MAJOR FINDINGS OF PNEUMATIC CYLINDER INVESTIGATION 265 7.12 RECOMMENDATIONS FROM PNEUMATIC CYLINDER INVESTIGATION 265
CHAPTER 8 EXTENDING FEEDER UNIT LIFE
8.1 SUMMARY 285
8.2 INTRODUCTION 286 8.3 CAUSES OF FEEDER FAILURES 297
8.3.1 Ore Leaks 288
8.3.2 Pointed and Short Plungers 290 8.3.3 Arcing 290 8.3.4 Fatigue of Piston Rods and Spools 297
8.3.5 Cylinder Bypass 291
8.3.6 Pot Off Line 292 8.4 ROD SEALS 292
8.4.1 Causes ofRod Seal Failure 292 8.4.2 High Temperature Exposure Control 294 8.4.2:1 Pot and Feeder Design 294 8.4.2:2 Normal Operation 295 8.4.2:3 Hot Pots 296 8.4.2:4 New Pots 297 8.4.2:5 Rod Seal/Bush Orientation 297 8.4.3 Speed. 298
8.4.4 Eccentricity - 298 8.4.5 Rod Seal Material 301
8.4.6 Rod Seal Design 303
8.4.7 Rod Wiper Design 306
8.4.8 Venting Rod Seal Air Leaks 307
8.5 PISTON SEALS 310
8.6 BROKEN PISTON RODS 310
8.6.1 Introduction 310
8.6.2 Fatigue Analysis 311 TABLE OF CONTENTS Xin
Page 8.6.3 Results and Conclusions of Fatigue Tests 314
8.7 ARCING 315
8.7.1 Insulation of the Assembly j/j
8.7.2 Cylinder Insulation J/7
8.7.3 Other Methods to Control Arcing 3 is
8.8 DOSING UNIT FAILURES 319
8.8.1 Broken and Worn Spools 319 8.8.1:1 Integrated Feeders 319 8.8.1:2 Independent Feeders 320 8.8.2 Spool Jamming 321
8.8.3 Spring Failure 323
8.9 MAJOR FINDINGS OF EXTENDING FEEDER UNIT LIFE 326
8.10 RECOMMENDATIONS OF EXTENDING FEEDER UNIT LIFE 326
CHAPTER 9 COST REDUCTION
9.1 SUMMARY 327 9.2 INTRODUCTION 329
9.3 PLUNGER SAVINGS 330 9.4 AIR SAVINGS 337
9.4.1 Ore leaks and Air Leaks 337 9.4.2 Air Leak Test Method 339
9.4.3 Boyne Feeder Maintenance Strategy 340
9.4.4 Non-routine Cylinder Overhaul Strategy 342
9.5 SAVINGS FROM REUSED FEEDERS 345
9.6 FEEDER TRACKING SYSTEMS 346
9.6.1 The Purpose of a Feeder Tracking System 346
9.6.2 Portland Feeder Reporting System 348
9.6.3 Tracking Systems and Quality Control. 350
9.7 MAJOR FINDINGS FROM COST REDUCTION 353
9.8 RECOMMENDATIONS FROM COST REDUCTION 353
CHAPTER 10 FEEDER DEVELOPMENT
10.1 SUMMARY 354
10.2 PROJECT HISTORY 355 TABLE OF CONTENTS XTV
Page
10.3 A2 FEEDERS, DRIBBLE FEED AND PULSE CHUTE 358
10.4 MULTIPLE DCVS PER POT OPTIONS 359 10.4.1 Split Header 360 10.4.2 Sequential Feed. 360
10.5 A3 FEEDER 362 10.5.1 A3 Pot Performance 362 10.5.2 Possible Use in Other Smelters 363 10.5.3 A3 Design Features 365
10.6 END STROKE SENSING 367
10.7 VALVE AND CYLINDER DESIGN 369
10.8 INTEGRATED VERSUS INDEPENDENT FEEDERS 370
CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS
11.1 SUMMARY 373
11.2 INTRODUCTION 374
11.3 PORTLAND RESULTS 374 11.3.1 Plant Performance 374 11.3.2 Improvements for Health 379 11.3.3 Worker Involvement. 383
11.2 ACTIONS AT ALCOA SMELTERS 384
11.3 ACTIONS AT NON-ALCOA SMELTERS 390
CHAPTER 12 SUMMARY OF RESULTS
12.1 RESEARCH OBJECTIVES 391 12.1.1 Feeder and Pot Performance 391 12.1.2 Feeder Development. 392 12.1.3 Conclusions on Achievements Against Objectives 393
12.2 FUTURE WORK 393
12.3 CONCLUDING REMARKS 394 INDEX OF FIGURES XV
INDEX OF FIGURES
Page CHAPTER 1
1-1 Portland Aluminium Smelter 1-2 Portland Aluminium Pot 3 1-3 Portland Pot Schematic 9 1-4 Potroom Layout 10 1-5 Aluminium Smelting Pots 12 1-6 Relationship between Pot Resistance and Bath Alumina Concentration 13 1-7 Early Pot Feeding System 18 1-8 Integrated and Independent Feeders 20 1-9 Integrated Feeder Operation (AEDD) 21 1-10 Crater below Integrated Feeder 22 1-11 AEDD Feeder Assembly 23 1-12 Pechiney AP18 Feeder 25 1-13 Pechiney AP18 Pot Showing Independent Feeder 26 1-14 A2 Feeder Operation 30 1-15 A3 Feeder Operation 31 1-16 Sequential Feed AEDD 32 1-17 Aeration Feeders 35 1-18 Pulse Chute Operation 37 1-19 Plunger End Stroke Sensing 41 1-20 Pneumatic Cylinder Parts 46 1-21 Portland Pot Feeder System 51 1-22 Portland Feeder Designs 57
CHAPTER 2
2-1 Off-Site Shot Size Test Rig 63 2-2 Portland AEDD Shot Size Graphs 66 2-3 Spool Insert Assembly 68 2-4 Parker Standard Checks 71 INDEX OF FIGURES XVI
Page 2-5 Effect of Cylinder Design on Portland AEDD Shot Size 74 2-6 Portland AEDD Repeatability Ex Plant. 76 2-7 Effect of Insert Stroke on Portland AEDD Shot Size 77 2-8 T6 Feeder Shot Size 79 2-9 Til Feeder Shot Size 81 2-10 Til #2 Feeder With and Without Insert 82 2-11 A2 Feeder Shot Size 83 2-12 Effect of Temperature on Springs 90 2-13 Cylinder Cushioning Operation 92 2-14 Effect of Cushioning on Portland AEDD Shot Size 95 2-15 Effect of Spool Design and Seal Age 97 2-16 Spool Wear Points 101 2-17 A2 Feeder with Double Outlet Chute 114
CHAPTER 3
3-1 Plant Crustbreaker Kinetic Energy and Pressure Comparison 126 3-2 Portland Feeder Mufflers 138
CHAPTER 4
4-1 Pot Pressure Sensing Points 145 4-2 Stroke Testing Unit 146 4-3 Pressure Traces - Old versus New Feeder 148 4-4 Pressure Traces - Speed Changes 156 4-5 Pressure Traces - On-Site and Off-Site Testing 151 4-6 Piston Rod Failure History 160 4-7 Pressure Traces-Number of Feeders per DC V On-Site 164
CHAPTER 5
5-1 Typical Worn Plungers 170 5-2 Trial Plungers 172 5-3 Plunger Diameter versus Life 178 5-4 Cast Iron Plunger Wear by Batch Number 183 5-5 Pot-to-Pot Plunger Wear Variation I88 5-6 Within-Pot Plunger Wear Variability 19° INDEX OF FIGURES XVII
Page 5-7 Cast Iron Plunger Wear Rate versus Wet Time per Crustbreak 195 5-8 Cast Iron Plunger Wear Rate versus Crustbreak Frequency 197 5-9 Effect of Crustbreak Frequency Change on Plunger Wear Rate 198 5-10 Plunger Wear by Month Installed 200 5-11 Effect of Cumulative Wet Time on Cast Iron Plunger Wear Rate 201 5-12 Plunger Wear Ratios 204 5-13 Calculated Plunger Life 205
CHAPTER 6
6-1 Dagged Plunger 231 6-2 Plunger Temperatures for 945°C Pots by Location 238 6-3 Plunger Temperature Profiles 248
CHAPTER 7
7-1 Cylinder Trial Equipment 257 7-2 Cylinder Trial Rig 258 7-3 Performance of Test Cylinders in Cylinder Trial 263 7-4 Cylinder 22A Rod Seal Leak Rate 269 7-5 Effect of Amount of Grease on Rod Seal Life 281
CHAPTER 8
8-1 Cylinder Rod Seal and Bush Designs 293 8-2 Seal Material Friction Coefficient 302 8-3 Weibull Analysis for Parker Rod Seals 304 8-4 Boomerang Brackets and Flange Insulation 308 8-5 Mounting Bolt and Bracket 309 8-6 Sketch of Broken Piston Rods 312 8-7 Weibull Analysis by Quarter of Piston Rods 314 8-8 Frequency of Jammed Objects 316 8-9 Weibull Analysis of Springs 325
CHAPTER 9
9-1 Bimetal Plunger 331 9-2 Calculated Plunger Cost per Tip-Day at A$75/Repair 335 9-3 Calculated Plunger Cost per Tip-Day at A$600/Repair 336 INDEX OF FIGURES XVHI
Page 9-4 Ore Leak History 339 9-5 Cylinder Overhaul Strategy Costs 344
CHAPTER 10
10-1 Multiple DCV's per Pot Feed Options 361 10-2 Independent Feed Option 361 10-3 A3 and Sequential Feeders 364
CHAPTER 11
11-1 Portland Development Implementation 375 11-2 Portland Feeder Changeout History 378 11-3 Life of Portland Feeders by Month Installed 380 11-4 Portland Feeder Performance - Summary 381 11-5 Alcoa Pot Feeder Performance-By Plant 382 11-6 Alcoa Feeder Performance - Summary 387 INDEX OF TABLES XLX
INDEX OF TABLES
Page CHAPTER 1
1-1 Smelters of the World 2 l-II NonAlcoa Point Feeder Systems 4 1-IJJ Duties of a Pot Feeder 5 1-IV 1991 Alcoa Feeder Costs (A$) 6 1-V Causes of Integrated Feeder Removal (1991) 16 1-VI Smelters Contacted in this Research 48 1-VTI Karmoy Cylinder Repair History 49 1-Vm Alcoa Feeder Recommendations 59
CHAPTER 2
2-1 Summary Results of Off.Site Shot Size Tests 73 2-II A3/AEDD Shot Size Comparison (Room Temperature) 84 2-UJ Anode Effects for Pots with/without Spool Inserts 86 2-IV Spring Force Comparison 98 2-V Shot Size Testing on New Pots 103 2-VI Summary of Shot Size Tests on an Operating Pot 105 2-VII Repeat Shot Size Runs using Same Feeder on an Operating Pot 106 2-VM Shot Size Comparison of Dwell Time and Speed on Feeders in an Operating Pot 107 2-LX A3/AEDD Shot Size Comparison (Operating Temperature) 109 2-X Comparison of Off Site versus Operating Pot Shot Size Tests for the Same Feeder 110 2-XI Things that Affect Shot Size 116
CHAPTER 3
3-1 Techniques to Break a Hole in the Pot Crust 120 3-U Plant Crustbreaker Kinetic Energy and Pressure Comparison 125 3-ID. Timeto Break a Hole at Zero Velocity 129 3-IV Causes of Blocked Feeder Holes at Portland 131 INDEX OF TABLES XX Page CHAPTER 4
4-1 Effect of Muffler and Speed Change-Location 1 154 4-U Effect of Muffler and Speed Change-Location 5 155 4-111 Dual Dwell Wet Times 162
CHAPTER 5
5-1 Parameters Potentially Affecting Plunger Wear 176 5-II Plunger Harness Trial Data 180 5-ffl Plunger Hardness Trial Data 182 5-IV Cast Iron Wear Rate Matrix 186 5-V Effect of Location on Plunger Wear 191 5-VI Plant Penetration Comparison 202 5-VTI Plunger Temperature Survey 207 5-Vni Effect of Plunger Speed and Dwell Time on Plunger Wear Rate 212 5-rX Plunger Material Comparison 214 5-X Effect of Tip Shape on Plunger Wear Rate 215 5-XI Kaiser Plunger Wear Tests 218 5-XJJ Plunger Materials Tested at Portland 221 5-XIH Effects of Plunger Wear Parameters for Cast Iron 227
CHAPTER 6
6-1 Dagged Feeder Survey Summary 234 6-II Link between Dags and Anode Effects 237 6-m Plunger Temperature Over a Three Week Period 241 6-IV Survey of Feeder Holes Open on Operating Pots 243 6-V Plunger Temperatures Relative to Chute Temperature 245
CHAPTER 7
7-1 Portland/Parker Cylinder Trial Data 262 7-U Portland/Parker Cylinder Trial Data Summary 271 7-III Recommendations from Cylinder Trial 272 7-IV Lubricant Decision Analysis Summary 275 7-V Final Lubricant Ratings 278 INDEX OF TABLES XXI
Page
CHAPTER 8
8-1 Causes of Failure of Rod Seals at Portland 292 8-II Maximum Temperature Limits for Seal Materials 296 8-ni Plunger Side Deflection 300 8-IV AEDD Spring Forces 324
CHAPTER 9
9-1 Portland Plunger Comparison 333
CHAPTER 10
10-1 Pot Feed Alternatives 356 10-H Feeder Trial Data 357 10-m Advantages and Disadvantages of Independent Feeders 371
CHAPTER 11
11-1 Portland versus other Feeders 376 11-11 Comments from Alcoa Plants on Feeder Development 386 11-111 Actions at Non-Alcoa Plants 390 INDEX OF APPENDICES XX11
INDEX OF APPENDICES
Appendix 1 Examples of Shot Size Control Charts
Appendix 2 Off.site Shot Size Test Data
Appendix 3 Alumina Sizing and Density
Appendix 4 Bob Seymour Memo on Feeder Recommendations
Appendix 5 Dick Talyor Memo on Results of Feeder Research
Appendix 6 A2 U.S.A. Patent
Appendix 7 Pulse Chute Patent Application
Appendix 8 A3 U.S.A. Patent
Appendix 9 Diary of Changes
Appendix 10 Plant Codes GLOSSARY PAGEXXIII
GLOSSARY
A2 Integrated feeder which feeds when the plunger is up; invented by the author. A3 Independent feeder invented by the author. AEDD Alcoa integrated feeder designed by the Alcoa Equipment Development Division. AEPD Average "anode effect per day" per pot. Airbum Combustion of an anode due to exposure of the surface to air when at elevated temperature. Alumina Aluminium oxide used for aluminium production. Anode Positive electrode of carbon.based material suspended from the superstructure into the molten bath. Conducting rods of copper or aluminium join the superstructure to the carbon. Anode Cover Material that covers the anodes to prevent air burning the carbon and to maintain heat balance in the pot. Anode Effect (AE) When a rise in voltage occurs on a pot due to the depletion of alumina in the bath to a level which the current can not be carried by the reduction of alumina and reduction of the electrode commences. Assembly The dosing unit of a pot feeder. Auxiliary Power In a smelter, power used which is not related to manufacturing of metal. Generally refers to power used in compressors, lighting and computers. Bath Liquid electrolyte used in a pot. Comprised mainly of cryolite but with low concentrations of alumina and calcium. Beehive Another name for a dag. GLOSSARY PAGE XXIV
BHN Brinell Hardness Number. Unit of hardness used for plunger wear assessment.
Blocked Feeders Blocked feeder hole in a pot causing buildup of alumina from the feeder. Also called mounding. Bridge Horizontal aluminium bar on the superstructure that suspends the rods of the anodes. Carbon Another name for anodes.
Cathode The negative electrode forming the bottom of the pot. Made of carbon-based material. Chisel Another name for a plunger.
(In) Control In statistical control. All values are within predicted probability limits...the process is predictable. Control chart A graph of values and a graph of variability between values with horizontal lines showing acceptance limits for statistical control.
Crucible Cylindrical vessel used to retrieve aluminium from the pot. May contain over 20 tonne of molten metal which is withdrawn from pots (tapped) by vacuum. Crust Concrete-like hard layer above and between anodes in a pot. Made of alumina and/or crushed bath. Crustbreakers The pneumatic cylinder, shaft and plunger unit that creates a hole in the crust for alumina to enter the bath. CE (Current Efficiency) Percentage of efficiency of metal production from a given amount of electricity relative to the theoretical conversion
rate. Cylinder bypass Piston seal leaks causing air to bypass through the cylinder.
Dag Buildup of bath material on the end of a plunger. DCV Directional Control Valve which supplies air from an air supply generally to an air cylinder. Activated by an electrical signal, an internal spool or poppet has its GLOSSARY PAGE XXV
direction changed by a solenoid when air needs to be applied to the inlet ports and the outlet ports need to be vented to the exhaust.
Design of Experiment Statistical technique to trial multiple parameters in a disciplined manner to identify causes of variability. Dwell Time Time from DCV movement when feeder (or crust breaker) starts to move down, to when the DCV moves to raise the feeder (or crust breaker).
Dense Phase Pneumatic conveying system supplying alumina to pot superstructure. Dosing Unit The component of a pot feeder that measures and delivers the shot of alumina. Also called the spool in the feeding assembly. Dough Ball Another name for a dag. Downstroke Movement of the piston vertically down a cylinder installed in a pot. Duct Fume duct. Endothermic Reaction which requires energy to occur. Energisers Spring or "O" ring type material that forces the seal against the moving part. Generally positioned between the seal and the fixed housing. FEC Front end cartridge. Rod seal unit on a pneumatic cylinder. Includes rod seal, rod bush, and wiper. Feeder Alumina feeding unit. This document refers to a feeder as the complete feeder/crustbreaking unit, but some plants (especially those which use independent feeders) may refer to a feeder as only the small pneumatic cylinder, shaft and dosing unit that doses the alumina.
Foot Another name for a plunger. GLOSSARY PAGE XXVI
Fume Duct Cavity in superstructure for pot gases to vent. May also refer to pipe from superstructure which transfers the gases to the scrubbing systems. Greens Newly manufactured anodes. Hornet's Nest Another name for a dag. HR Cast Iron Heat Resistant cast iron developed by Alcoa. ICC Team In Control and Capable Team - co-ordinators of research and implementation of improvements in Alcoa plants. ID. Internal diameter.
Incoloy Trade name of a high chrome/nickel stainless steel alloy used for plungers Inconel Trade name of a high chrome/nickel stainless steel alloy used for plungers Independent Feeder Crustbreak and alumina dosing are operated by separate pneumatic cylinders. Integrated Feeder Crustbreak and alumina dosing are mutually operated by one pneumatic cylinder.
Kidney Plate Valve on superstructure that isolates alumina entry into the feeder.
LCL Lower Control Limit - usually 2.5% of values fall below this limit.
Lights Colloquial expression for an anode effect. Refers to illumination of a light bulb mounted on a pot when the voltage rises above 6 volts during anode effect.
Location Position of feeders in the pot superstructure - #1 closest to narrow (control panel) aisle and #5 closest to wide (tap)
aisle.
Mack Valve Colloquial expression used for a directional control valve (DCV). Is a trade name for a DCV.
Metal Pad Layer of molten aluminium at the base of a pot. Mounding Blocked feeder hole. GLOSSARY PAGEXXVII
Muffler Exhaust silencer on a DCV. Nicrofer Trade name of a high chrome/nickel stainless steel alloy used for plungers O.D Outer diameter. Oiler/Filter Unit mounted upstream of a DCV. Ore Aluminium industry jargon for alumina. Ore leak Unmetered alumina leak into the pot from the feeder or kidney plate. "O" ring Ring shaped seal used often as an energiser behind a rod or piston seal.
Outliers Extreme values. PA Spec Cast iron specification plunger developed in this study for Portland Aluminium.
Pareto Chart Bar chart with values decreasing from left to right Penetration Distance that the plunger is wetted by molten bath. Pick Another name for a plunger. Plugger Another name for a plunger or feeder/crustbreaker unit. Plunger Metal chisel used to break a hole in the crust above the liquid components of a pot.
Plunger wear rate Reduction in O.D. of the bottom edge of the plunger over time.
Pot Electrolytic cell used for manufacture of aluminium. Pot dressing Covering anodes in operating pots using a rake type device; usually not related to anode setting.
Potline An array of pots electrically connected in series. Potlining Carbon material that lines a pot to protect the insulation from corrosion.
Potroom Half of a potline.
Prebake pots Pots that have anodes baked prior to position in the pots; alternative to Soderberg pots where the anode is formed
from a paste in the pot itself. GLOSSARY PAGEXXVIH
PTFE Polytetrafluoroethylene; often called by the trade name Teflon. Reactors Systems to scrub fluorides from the pot vent gases. Restrictor A 12mm orifice plate in the air lines to and from a Portland pot. Used to reduce stroke speed of the feeder cylinder piston.
Rod The vertical bar attached to the carbon block of an anode. This is suspended off the bridge in the pot. Scheduled AE Anode effect which is programmed (scheduled) to occur by manual intervention. Set Installation of new anodes and removal of spent anodes. Sequential Feed Each feeder in the pot operates at different times. For example, order of feeding may be 3,1,5,2,4 in sequence for a 5 feeder pot.
S.D Standard deviation. Sigma One standard deviation of a sample. Silencer Straight.through silencer on the DCV of the pot feeder air supply system. Similar principle to arifle silencer .
Spa Calcium fluoride used to control the hardness of the crust above the molten electrolyte in the pot. Lower spa gives harder crust.
Spool The measuring unit that measures and delivers the shot of alumina. Also known as the dosing unit.
Statistical SignificanceWher e there is less than (usually) 95% that something occurred by chance and chance alone.
Stroke The movement of the piston in a pneumatic or hydraulic cylinder from one end of the barrel to the other
Super Superstructure. Superstructure Central storage vessel above the pot cavity. Holds alumina, feeders and anodes. Fume duct vents from
superstructure. GLOSSARY PAGEXXLX
Tap Removal of metal from a pot via suction. Tapout Pot leaks molten metal or bath and has to be cut out of the process. Teflon A trade name of PTFE; a high temperature slow wearing plastic commonly used for seals on pneumatic cylinders. Thermoset Hardenning of a material due to heat. Commonly found for rubber "O" ring seals used as energisers on pneumatic cylinders under extreme temperature excursions. UCL Upper Control Limit - usually 2.5% of values fall above this limit. Upstroke Vertical movement of the piston from the low point. Viton A trade name for a high temperature rubber, like hydrocarbon material commonly used for seals on pneumatic cylinders. Wallbox Control cabinet on the narrow aisle side of each pot to allow manual intervention of pot control, and to monitor pot performance. Weibull Analysis Logarithmic graph of failure data that is used to predict failure rates and mechanism of failure. The Y axis shows percent failed and the X axis shows life of the component or part. A curve fit is conducted and the two critical parameters that explain the curve are displayed viz. (3 (shape parameter) and TJ (age when 62.3% have failed). When P is 1.0, failures follow an exponential relationship, and when 3.5 the distribution is normal. In most of the failures in this text, p is close to 1.0. (Refer Juran (96)
page 23-34 for more details).
Wet Time Time that the plunger is fully extended and potentially immersed in molten bath.
Wiper Outer unit of the rod seal (front end cartridge) that wipes material off the piston rod of pneumatic cylinders. FIGURE 1 -1 PORTLAND ALUMINIUM SMELTER
Photo courtesy Portland Aluminium, Portland, Australia CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGEI
CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY
1.1 WHY ALUMINA FEEDER TECHNOLOGY IS IMPORTANT World production of aluminium is about 20 million tonnes per annum (Table 1-1). Alcoa is the largest manufacturer of aluminium in the western world with interests in the USA, Brazil, Europe and Australia. Their newest smelter is Portland Aluminium at Portland in western Victoria, Australia which was commissioned in 1986 (Figure 1-1).
A key factor in maximising the efficiency of smelting alumina to aluminium is the control of the concentration of the alumina in the molten electrolyte of the electrolytic cell (commonly called a "pot") (Figure 1-2). The concentration needs to be kept within a tight tolerance in the 2% to 5% range or the pot will rise in voltage (if too low) or form sludge (if too high). The ideal range is 2-2.5%, though many pots operate closer to 3-5%.
Low concentration causes starving of the pot. This is called an "anode effect". The resulting high voltage causes a rapid rise in temperature which affects the heat balance. Sludge (commonly called "muck") drops to the bottom of the pot, increases voltage across the pot and the pot increases temperature rapidly. The closer the pot operates to an anode effect, the more efficient and cost competitive the process but the more sensitive it is to upsets. One can see that the pot operates on a knife edge of concentration. Either side of this acceptable range, efficiency drops and profit margins reduce.
The high voltages associated with low alumina concentrations are caused by the generation of CF4 and C2F6 which are "green house" gases which may contribute to global warning. Research work has shown that CF4 and C2F6 have about 5,000 and CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 2
10,000 times respectively more global warming capacity than carbon dioxide (1). For these reasons, for virtually all smelters, a major target is to reduce anode effects. Large smelters use 500-700MW of power for production rates up to 500,000tpa. The efficiency of the pot is the main contributor to this power demand as potlines and pots consume 12-14kWhr/kg of metal produced which comprises up to 80% of smelter power usage.
TABLE 1-1
SMELTERS OF THE WORLD
No. of Smelters Capacity 100% Share (tpa x1000)
Russia (Government) 11 0 3008 Alcoa 8 6 1778 Alcan Smelter & Chemicals 13 7 1735 Reynolds Metals 4 5 1018 China (Government) 17 1 898 Am ax 1 4 773 Aluminium Pechiney (Government) 4 6 723 Venezuela (Government) 0 2 541 Tadzhikistan (Government) 1 0 520 Vereinigte Industrie Unternehmungen A 3 3 499 Maxxam Group Inc. 0 4 445 Alusuisse-Lonza Holding Ltd 2 3 349 Bahrain (Government) 0 1 344 Instituto Nacional De Industria (Gvmnt.) 0 3 336 Norway (Government) 0 5 334 Conzinc Riotinto Of Australia Ltd 0 3 275 Romania (Government) 1 0 270 Ohio River Associates 1 0 245 Dubai (Government) 1 0 244 Votorantim 1 0 240 Companhia Vale Doe Rio Doce 0 2 224 Hodgovens 2 1 220 EFIM-MCS 3 1 218 Noranda Mines Ltd 1 0 215 Austrian Industries AG (Government) 1 3 209 India (Government) 2 0 209 Kazakhstan (Government) 1 0 200 Others - ~ 4610 20340
Notes: (i) Data from Pawlek (94) updated in 1994. (ii) "Capacity" is installed capacity but not necessarily at the level that the smelter is currently operating. (iii) "100%" refers to the smelters owned by the company and "Shared" refers to the number of smelters that the company has shares in. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 3
FIGURE 1 - 2
PORTLAND ALUMINIUM POT
Photo coutesy Portland Aluminium, Portland, Australia
To achieve optimum concentration in the cell, the control of alumina feed to the pot is critical. Almost all of this material is delivered by the alumina feeder.
Each pot has feeding equipment that varies in design depending on the pot design. Modern plants use 1-5 point feeders that dose about l-2kg of alumina into the pot about every 2-4 minutes. They are called "point" feeders as they deliver the dose into a CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 4 localised point, rather than on a large slot in the crust above the electrolyte. Alcoa plants have 1-3 feeders per pot except for Portland which has 5 feeders per pot...the largest number of point feeders in any pot in the world. Table l-II shows some feeder systems for non-Alcoa plants.
TABLE 1-D
NON-ALCOA POINT FEEDER SYSTEMS
INSTALLED FEEDERS/ O.D. STROKE ' SMELTER CYLINDER VALVES FEEDERS POT (mm) (mm) PLUNGER
D1 640 2 Atlas Copco 200 600 Stainless Atlas Copco 2 Mechman 200 600 Stainless Atlas Copco D2 888 4 Atlas Copco 200 500 AISI 309 Atlas Copco 4 CPOAC 200 500 AISI309 Atlas Copco D3 128 2 Atlas Copco 160 600 Alloy Casting Atlas Copco D4 - 4 - 200 600 - - D6 960 4 Atlas Copco 200 500 Z12CN Atlas Copco D8 - 4 - 200 600 - - D9 1920 4 Atlas Copco 200 500 Nicrofer Atlas Copco 4 Atlas Copco 200 500 Hastelloy X Atlas Copco D10 1920 4 50%CPOAC 200 500 31 OSS Parker 4 50%Atlas Copco 200 500 31 OSS Parker D13 684 3/2 Atlas Copco 160 - Cast Iron Atlas Copco D14 720 4 - 200 600 NS24 - D15 - 4 - 200 600 31 OSS - D16 1152 4 Atlas Copco 160 500 31 OSS Atlas Copco D17 - 4 Atlas Copco 200 - - - D18 3600 4 Atlas Copco 160 - - Atlas Copco D19 1056 4 50%Atlas Copco 200 - 309SS Atlas Copco 4 50% CPOAC 200 - 309SS - D20 - 4 Atlas Copco 160 - - Atlas Copco D21 - 4 Atlas Copco 160 - - Atlas Copco T2 352 2 Atlas Copco 125 508 Cast Iron Ross T3 328 2 Mechman 125 500 Alloycoating Mechman T4 864 3 Lindberg 125 356 304SS Mac T8 1920 4 Parker 125 470 Steel/304SS Parker T5 768 3 Lindberg 100 356 304SS Mac T14 768 2 Lindberg 100 356 304SS Mac T10 240 2 Terry 125 356 Cast Iron Ross T17 1080 3 50% Scheffer 100 356 Cast Iron Mac Notes: (i) "T" refers to integrated and "D" refers to independent feeder plants. (ii)"-" refers to data not being available.
The capital and operating costs for feeders are a significant cost to smelters as there are up to 3,000 feeders in large modern plants at capital costs of up to $30,000/pot for several hundred pots. The duty is arduous due to the high temperatures (up to 1,100 C), corrosive liquids and gases (cryolite, HF, H2S, SO2) abrasive crust above the electrolyte, and the need to provide safe and hygienic conditions for personnel in operating and maintaining the equipment. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 5 Pot feeder designs used in the modern smelters have basic requirements to:
(i) measure an accurate and repeatable dose. (ii) deliver the dose every time it is required. (iii) ensure the dose enters and mixes efficiently in the molten electrolyte (bath).
If these actions are carried out effectively, low frequency of anode effects should result. Table I-HI lists the main duties of a pot feeder. Note that capital/maintenance costs are only one of the major aims of a good feeder. Table 1-IV shows typical maintenance costs. These costs are not trivial, however, it is generally considered that these costs are small in comparison to the greater needs of control of anode effects and pot efficiency. If market prices of aluminium are low, these costs are important factors in the cost competitiveness of a smelter.
TABLE l-III DUTIES OF A POT FEEDER
ACCURATE MEASURE AN ACCURATE AND REPEATABLE DOSE (SHOT SIZE)
DOSE IS DELIVERED ENSURE ALL THE DOSE ENTERS AND MIXES EFFICIENTLY IN THE MOLTEN ELECTROLYTE (BATH)
" RELIABLE DELIVER THE COMPLETE DOSE EVERY TIME REQUESTED
SAFE SAFE AND HYGIENIC OPERATION
COST EFFECTIVE MINIMUM OVERALL COST...CAPITAL AND MAINTENANCE
The effect of the design on the people who operate and maintain the pot feeders needs special attention. Changing feeders in an operating pot results in greater exposure of people to falls (up to 3 metres), heat, dust, and fumes as they operate the cell. Even maintaining the feeders creates hazards due to manual handling of heavy equipment and potential exposure to fluorinated alumina. This is an area of responsibility that CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 6 engineering must address in the development of any piece of equipment. This facet of engineering design is often significantly lacking.
TABLE 1-IV 1991 ALCOA FEEDER COSTS (A$)
PLANT # FEEDERS # FEEDERS AVE LIFE TOTAL $PER INSTALLED CHANGED (months) COSTS CHANGE (A) (B) (A/Bx12) ($ pax 1000) PORTLAND 2040 2400 10 1523 635 T6 734 700 12 446 635 T7 1824 4794 5 2492 N/A T9 1548 940 20 908 966 T11 564 200 34 148 738 T13 496 400 15 406 1015 T15 1936 940 24 908 966 T16 1700 880 23 457 520 D12 984 694 17 748 1077 ALL 11826 11948 12 8035 718
Notes: (i) '*' indicates this is an estimate. (ii) "average $/change" does not include D12 as this is the only plant which is not an integrated feeder plant. (iii) Assumed A$ = US$0.65 (iv)" T" refers to integrated and "D" to independent feeders.
It is desirable, therefore, to optimise the efficiency of aluminium smelting cells for
several major reasons:
(i) potential global warming effect (ii) cost control
(iii)occupational hygiene/safety. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 7
1.2 LITERATURE SURVEY
1.2.1 Aluminium Smelting Technology
Aluminium is the third most plentiful metallic element in the earth's crust (7.3%) (5). It is found in 250 oxides exceeded only by oxygen and silicon. One of the oxide hydrates (bauxite) is the dominating raw material for the production of aluminium. Extraction of the metal has three main stages:
(i) mining of the ore (bauxite); (ii) refining the ore to recover alumina (aluminium oxide); and (iii) smelting the alumina to produce aluminium.
The refinery Bayer process is the prevalent method of refining bauxite to make alumina. In this process, a caustic soda solution is used to dissolve the ore and alumina crystallises as a 100 micron solid with generally a small particle size range. (Descriptions of the mining and refinery processes are explained by Prider and Grojotheim et al. in references 5, 6 and 7.)
The smelting process comprises the electrolytic reduction of alumina (A1203) in a bath of molten cryolite (sodium aluminium fluoride, Na3AlF6) using carbon anodes:
2A1203 + 3C + electricity = 4A1 + 3C02
As the current goes through the pot, electrons exchange between the disassociated alumina in solution to produce aluminium:
Al+++ + 3e' = Al
Electrical current flows through the pot via anodes, through molten liquids of bath and metal, and out the carbon cathode base of the pot via a matrix of aluminium bus bars to the next pot where the process continues in a similar manner. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 8
Figure 1-3 shows the main features of the Portland pot which is typical of normal prebake pots (though considerably larger; 1.5-3 times the size of most prebake pots). A pot is basically a large tank which has a 100-3 00mm layer of molten aluminium below a 100-300mm layer of molten electrolyte (bath). A rectangular silo of alumina (the superstructure) is centrally mounted for alumina feeders to dose alumina into the bath at regular intervals. Carbon anodes are suspended on a large horizontal busbar (the bridge) off the superstructure. These carbon blocks (of up to one tonne in size) have either aluminium or copper vertical bars (rods) that are fixed to their top surface usually with cast iron or pressed carbon material. The vertical "rods" are attached to the bridge which moves up and down by electric motors in response to the voltage across the pot using automatic computer programs.
Up to about 260 pots can be in a potline using the same current path (Figure 1-4). The potlines are generally divided into two potrooms of equal length, though sometimes (as in Figure l-4(a)) this may be under the same roof. Despite being of only half the length of a potline, potrooms can be over a kilometre in length. Another feature of potlines is the use of overhead cranes (one of which can be seen in Figure l-4(b). These are used to service the pots and are used to change anodes (setting) and extracting metal (tapping). In many older potlines, cranes also are used to supply alumina to the pots by hoppers contained in (or hung off) the cranes. In more recent potlines pneumatic conveying or air fed systems are used instead of cranes for alumina supply of pots to reduce the amount of spilled alumina and to reduce the usage of cranes. Retrofitting dense phase systems is becoming more common since the mid 1980's.
Aluminium smelting is a continuous process with the alumina being dissolved in the bath. The aluminium (being denser than the 950-960°C bath) forms in the bottom of the pots. Alumina concentration is controlled at 2-5%. Molten metal is tapped on a 1-2
day routine by use of vacuum tanks (called crucibles). Gases (mainly C02, S02 and HF) are released from the solution. n
>
Q § P § CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 10
FIGURE 1 - 4 POTROOM LAYOUT
(a) End to end prebake (left) and Soderberg(right) pots except for 3 pots on lower left which are prebake pots).
(b) Side by side prebake pots
Photos courtesy GA Metall AB, Granges, Sweden
The pots are arranged in a series of cells forming the potline, with each cell operating at 4-5 volts DC at a current of 60-3 OOkA. Direct current passes from carbon anodes through the bath to the cathode of the cell and then to the anodes of the next pot (and so on). Steel bars embedded in the cathode carry the current out of the pot while the pots themselves are connected via an aluminium bus-bar system.
Generally there is a similar current per unit area (current density) of pots, so it is usual to refer to the relative size of a pot by referring to the current applied to a potline. Most CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 11 potlines are 120kA to 200kA. The Aluminium Pechiney AP30 and the Alcoa model 817 pot (Portland) are the only potlines operating close to 3 OOkA. These are the largest pots operating at present and are about 10m long by 5m wide using anodes of about a tonne mass. This compares with the earlier pots of only a few kA and anodes so small that they could be carried by hand. There has been rapid growth of the industry since Charles Martin Hall and Paul Heroult invented the process in 1886.
These cells are of generally two different design types (Figure 1-5).
(a) Soderberg pots
Currently, 30% of aluminium production is from Soderberg pots (mostly from CIS countries) (8). These pots use a carbon based anode paste (which is a mixture of coke and pitch) that is dropped into a steel casing hanging above the pot. The carbon paste is baked in the pot itself by virtue of the heat coming out of the molten bath while vertical or angled steel pins (or studs) hold the anode and bring the current into it. This type of pot has poor fume control and generally poor efficiency. There are two types: vertical (as shown in Figure l-5a) and horizontal stub (which is the steel conducter in the anode paste).
(b) Prebake pots
These pots have the carbon paste baked prior to being placed in the pot. They have the carbon based anode compressed or vibrated into a block which is then baked at about 1,100°C to precondition the 'green' anodes prior to placement into the pot. This results in better pot fume control and the pots run at superior efficiency. These anodes are changed at about 20-30 days as the carbon burns away with the oxygen generated from the dissolution of the alumina. Carbon dioxide and other gases are driven off into collection fume systems which are then routed to scrubbers to collect any fluoride in the gas.
Soderberg pots were the main production units initially, but better designs of prebake pots have developed such that almost all new pots commissioned since the 1960's have CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 12 been prebake. A large number of potlines are now retrofitting prebake pots into Soderberg lines or converting the Soderberg pots to point feed.
FIGURE 1 - 5 ALUMINIUM SMELTING POTS
LIQUID PASTE
ALUMINA
BAKED ANODE FROZEN LEDGE
LIQUID ALUMINIUM yj £\JM1I
(a) Soderberg
ALUMINIUM RISER
FROZEN LEDGE
(b) Prebake
This document concentrates on crustbreaker/feeder units (generally called feeders) in prebake pots. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 13 1.2.2 Objectives of Alumina Feed Control
The control of the concentration of alumina in the bath is a critical factor in the efficiency of a pot. As the concentration drops, the voltage in the pot falls to a minimum at about 2.5% w/w alumina in solution. Once this point is reached, there is a very rapidrise i n voltage (Figure 1-6). This rise in voltage is called an "anode effect".
FIGURE 1 - 6
RELATIONSHIP BETWEEN POT RESISTANCE AND BATH ALUMINA CONCENTRATION
1.82 r
1.80 • -
o 1.78 - X m E 1.76 x: O UJ 1.74 - CoO CO LU 1.72 -
1.70 1 3 4 ALUMINA (%)
It is desirable to run at the least number of anode effects as is possible in a day (or days), as generation of anode effects creates rapid short term energy input into the pot. This upsets the heat balance, current efficiency and subsequent power costs. However, to run at too few anode effects, there is a risk that the pot has undissolved alumina (muck) in the bottom of the pot which causes uneven current flow, unstable metal level CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 14 and loss of current efficiency. A mucky pot can become very hot and dangerous to operating personnel and is also very inefficient. Muck is a far worse evil than too many anode effects, so smelters have traditionally programmed (scheduled) anode effects to ensure that the alumina concentration does not get too high and create muck.
Recent concern that the gases given off during anode effects are potentially harmful to the atmosphere, and the findings of a close relationship of low anode effects to better current efficiency has directed smelters to reduce anode effects yet still prevent muck from forming. This is a difficult task as pots are operating much closer to instability. To achieve these targets, better feeder efficiency and feeder life plus smart computer feed control are required.
Hence, it is important to control the alumina concentration within a narrow range. Feeding systems generally comprise 80% of alumina input to a pot (with the balance coming from the material that covers the anodes), so control of the feeding system is very important to controlling pot performance.
To dose alumina into the liquid bath, a path has to be broken into the hard crust formed over the top of the anodes. This rock-like crust is generally a mixture of alumina and solidified bath (called anode cover). The hard crust, even when broken, often reforms quickly, so frequent breakage is required. A maximum time between feeds of less than 10 minutes is used in practice and feeding about every 3 minutes being typical.
Another complexity is that the bath is highly corrosive and dissolves most metals rapidly. The abrasive nature of the anode cover and crust also tends to wear the plunger away. Hence, the crustbreaking device must be capable of resisting this attack yet not be uneconomic by being too expensive. Metals such as Inconel or high nickel based materials are generally efficient materials, but costs are very high and considered by many to be uneconomic. To make matters worse, the crustbreakers must withstand high temperature cycles (100°C to 1,000°C every 3 minutes) and hydrogen fluoride gas. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 15
The feeding device must have a long life to avoid exposure of people to heat, dust and fumes during change-out. Corrosion of plungers contributes to contamination of aluminium with iron or other elements in the material (especially if their life is short... 1-2 years). For potlines of several hundred pots with 2 to 5 feeders per pot, capital and maintenance costs quickly become significant if feeder life is low.
Table 1-V shows data on the causes of failure of feeders at 11 integrated feeder plants. This shows the variety of both the causes and the frequency of each type of failure despite all plants having basically the same feeder design. Chapter 9 discusses causes of failure and how to prevent them.
To get the most out of a pot horizontal area, the carbon area in the pot should be maximised so maximum current can pass through the pot, thus, reducing the number of pots and capital cost per tonne of aluminium. As this happens, less area is available for pot feeding. Also, there is a need to conserve energy as one of the highest costs of operating a pot is the cost of power; breaking the crust and leaving open holes loses energy, so is not desirable. Clearly, a balance of priorities is needed which results in the duties of a pot feeder being difficult to achieve (Table I-in).
Although many papers have been published on dissolution of alumina, almost all published literature on pot feeders or associated systems is contained in patents. A lot of work is done on a plant-by-plant basis but not published. Hence, other smelters often are not aware of developments in this competitive industry. No papers have been traced on the efficiency of operating feeders other than references in some papers by Pechiney which provide a summary of patents (17) or general papers on the benefits of potlines
(19, 20). CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 16
TABLE l-V CAUSES OF INTEGRATED FEEDER REMOVAL (1991)
AVERAGE BROKEN OPERATING ANNUAL ORE POINTED POT OFF SHORT BROKEN NOT SMELTERS LIFE ARCED PISTON FEEDERS OVERHAULS LEAK PLUNGER LINE PLUNGER SPOOL FEEDING (months) ROD PORTLAND 2040 2400 10 28 17 40 1 11 3 T2 352 168 25 95 5 T6 734 700 13 15 60 10 10 15 T7 1824 4794 5 60 10 20 10 T9 1548 940 20 10 85 5 T10 240 310 9 80 10 5 5 T11 564 200 34 20 40 35 5 T13 496 400 15 15 75 10 T15 1936 940 25 80 20 T16 1700 880 23 10 60 20 5 5 T17 1080 364 36 75 25 ALL 12514 12096 12 44 12 13 13 8 5 3 2
SUMMARY OF ALL 11 SMELTERS SURVEYED
BROKEN PISTON ROD S% NOT FEEDING 2%
BROKEN SPOOL
SHORT PLUNGER 8%
POT OFF LINE 13%
POINTED PLUNGER 12% CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 17
1.2.3 Initial Prebake Feeding Systems
The early feed systems were based on the feeding systems of the Soderberg potlines which often drove a large toothed wheel along the side of the pot and poured alumina onto the top of the holes (9 and 10). Figure l-7(a) illustrates one such design. This method was difficult to control the unmetered alumina supply and muck often developed which caused pot instability.
Because of the large number of existing Soderbergs (especially in the CIS), better feeders are being developed that have similar features to prebake point feeders (14), use of fluidised feeders (13), or fluidised airslides (14).
The next generation of feeding systems was a bar break (16,17). Figure l-7(b) illustrates one such design. Along the centre of the pot a large guillotine, several metres long, is used to break the crust. This device is operated by large air cylinders that, by necessity, have to be located close to the dust, heat and fumes. The feeding system is generally a dosing unit above the bar that discharges a dose of 30-70kg on top of the slot in the crust.
The main problem with the barbeak design is the haphazard nature of alumina entry into the bath. There is no guarantee that all the alumina reaches the bath and at what rate. Also, the large area of the bar posed problems on economic materials of construction for long life. Although better that the 'toothed wheel' of the Soderbergs, bar break dosed too large a slug of alumina into the pot and pot efficiency suffered. Reverdy (16) gives a good description of the patents of the Soderberg and barbreak feeders in early prebake pot designs. Many operating plants still use barbreak feeders, but some have converted or are converting to point feeders (e.g. Alusuisse (18) and Comalco respectively).
In 1970, Pechiney patented an improved device for pot feed based on crustbreakers and dosing into multiple drilled holes (15). The principle was in line with later point feeder development, but the high cost of the large device and the need to use overhead cranes resulted in the design not being pursued. CHAPTER l INTRODUCTION AND LITERATURE SURVEY PAGE 18
FIGURE 1 - 7 EARLY POT FEEDING SYSTEMS
Adapted from Patent SU1301875 (Ref 9)
(a) Sidebreak
Adapted from Patent NO-143506 (Ref 17) (b) Barbreak CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 19
The superior technology for alumina feeding is "point feeding" where an accurately measured small dose (under 5kg) of alumina is dosed into a small hole (3 0-120mm diameter) in the crust. The alumina is delivered into the pot via 1-5 dosing systems that incorporate a crustbreaker and a dosing unit.
There are two basic types of point feeders (Figure 1-8):
Integrated feeders...one air cylinder used for dosing and crustbreaking in a combined design; and
Independent feeders.separate cylinders for dosing and crustbreaking such that the two functions operate independently.
Both types of feeder have a crustbreaker tip - called a plunger - which breaks a hole in the hard crust above the molten bath to provide a path for the alumina to get into the bath. Plungers are called tips, pluggers, chisels or picks in various plants.
1.2.4 Integrated Feeders
A couple of designs were developed in the 1960's by Alcoa (Figure 1-9) and Reynolds (21). In the integrated C^OA^ the downward movement of the crustbreaker releases a premeasured volumetric dose onto the top of the plunger. The plunger is still in the hole when the alumina is released. When the plunger moves upwards, most of the dose falls into the hole. The alumina tends to build up a volcano shape around the broken holes for integrated feeders (Figure 1-10). The alumina dose has to slip down into the hole after the plunger lifts out of the way. Thus, the feed is only indirectly reaching the bath and is affected by the angle of the crater and blowback of gases out of the hole.
Integrated feeders generally use a 100mm to 150mm O.D. pneumatic cylinder which activates the crustbreaker and dosing system as it strokes. They generally incorporate high speeds to feed quickly the alumina out of the dosing unit and to develop high kinetic energy to break the crust. This design is relatively cheap, simple and easier to maintain than the barbreak, as the feeder can be lifted out of the pot by crane to repair CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 20
FIGURE 1 - 8
INTEGRATED AND INDEPENDENT FEEDERS
INTEGRATED INDEPENDENT CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 21
FIGURE 1 - 9
INTEGRATED FEEDER OPERATION (AEDD)
ALUMINA FROM SUPERSTRUCTURE
ALUMINA DOSE
ASSEMBLY SEALING FACE
DELIVERY OF DOSE CUP SEALING FACE
PLUNGER UP
% %
PLUNGER DOWN CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 22 any failures. This type of design became the norm in new smelting lines until the 1980's when independent feeders were developed.
FIGURE 1 -10
CRATER BELOW INTEGRATED FEEDER
Note: This crater could have been the result of a blocked feeder hole. The shape is similar to that of a normal feeder hole crater.
USSR Patents (24 and 25) appear to be similar to the Reynolds/Alcoa integrated feeders but incorporate baffles in the tank to push alumina into the dosing system. Baffles are an attraction if the alumina is badly segregated, however, the more correct action should be to minimise the segregation in the first place and avoid the extra capital cost of equipment for each feeder.
The most popular integrated feeder is the Alcoa AEDD developed by the Alcoa Equipment Development Division in the 1950s and 1960s (Figure 1-11) with the Comalco feeder at Boyne Smelters being of similar concept but with a different feeding mechanism. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 23
FIGURE 1-11
AEDD FEEDER ASSEMBLY
SUPERSTRUCTURE "TORPEDO TUBE" PLUNGER SHAPT -ASSEMBLY BODY SPRING VEN HOLES KIDNEY PLATE* BASE OP SUPERSTRUCTURE
ALUMINA DOSE^ INSERT
ILATION BAND
GUIDE
HARDENED OR BRONZE INSERT
SUPERSTRUCTURE
PLUNGER PIN-
PLUNGER CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 24 1.2.5 Independent Feeders
Independent feeders separate the function of crustbreaking and feeding though they are kept located in the same area of the pot. A large air cylinder (150-200mm O.D.) is used as a crustbreaker and a small cylinder (40-50mm O.D.) is used to dose alumina.
In any pot, it is not necessary to break a hole every feed as the crust takes minutes to hours to cover the hole. As opposed to the integrated feeder that can only dose when the crustbreak mechanism is extended, the independent feeder need only break a hole when it is deemed prudent. Plants can break a hole every 2-3 feeds, thus reducing the cycles of the cylinder and frequency of penetration of the plunger into the bath. Maintenance costs and improved feeder life are evidenced. Reverdy (26) maintains that another major advantage of independent feeders is direct feed into the hole and less air usage. (Reverdy is a Director of Aluminium Pechiney.)
The most common independent feeder is that based on Pechiney's patents of Gerphagnon and Wolter (22) and Bonney and Gerphagnon (23) (Figure 1-12 and 1-13). The design is used by Pechiney in their API8 (180kA) pots and with a revised model in their AP30 (3 OOkA) pots. Several other companies have adopted similar styles without infringing patent rights e.g. Hydro Aluminium, VAW, Reynolds, Granges Metall, Aluar. This type of feeder is rapidly catching up to Alcoa's integrated feeder in popularity and will soon be the most popular feeder after commissioning of the ALUSAF smelter in South Africa.
Alusuisse (Swiss Aluminium) recently have constructed high efficiency retrofit packages for feeding and dense phase supply of alumina to the pot (18). Some of their plants use an independent feeder that has a movable spout to direct the flow into the hole in the crust in the one mechanism, while others use a similar design to Pechiney. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 25
FIGURE 1 - 12
PECHINEY AP18 FEEDER
BOX FOR AIR CONNECTIONS (SEPARATE AIR VALVES ON POT CRUSTBREAKER FOR EACH CRUST CYLINDER BREAK & FEED CYLINDER)
SPRING (TO ENSURE DOSING SYSTEM MATES WHEN DROPPED INTO PLACE BY CRANE) FEED CYLINDER
AIR BLEED FROM DOSING CYLINDER
FIXED CHUTE IN FIXED CHUTE IN SUPERSTRUCTURE SUPERSTRUCTURE
PLUNGER SHAFT GUIDE FOR SHAFT OF DOSING UNIT
AIR BLEED FROM CRUSTBREAKER CYLINDER
BUSH HOUSING
"DAG" SCRAPER
FLAT BOTTOMED DELIVERY CHUTE PLUNGER
Adapted from Gerphagnon and Wolter Patent: USA 4,437,964 (22) CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 26
FIGURE 1 -13 PECHINEY AP18 POT SHOWING INDEPENDENT FEEDER
DOSING NOZZLE
DAG CLEANER
Adapted from thefront cove r of Light Metal Age, December 1992.
ALUMINA OVERFLOW GAP
Adapted from the front cover of Kongerud (74). CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 27
1.2.6 Comparison of Integrated and Independent Feeder Designs
The integrated feeder preheats alumina because the alumina dose rests on the hot crust before it slips into the hole, but the independent feeder allows direct entry of alumina into the hole. The latter may provide better mixing as the alumina has a vertical and horizontal velocity as it hits the bath. The alumina dose enters as a slug rather than a gradual flow; which reduces dust loss into the pot cavity (27). However, there is no pre heating of the alumina prior to entry into the hot bath which may cause the dose to drop to the bottom of the pot as a lump until it is broken up by the turbulence of the metal and/or it reaches dissolution temperature. This mechanism can result in muck forming if this does not happen effectively to all the dose.
Examination of Figures 1-10 and 1-13 illustrate the smaller cavity for independent feeders. Note that Figure 1-10 has a very large crater which is not typical of all integrated feeders, but it does show the general trend of crater shape. Also of note is the alumina overflow gap on Figure 1-13(b) which provides a path for alumina to still get into the pot if the dose chute is blocked. Integrated feeders do not have a small nozzle so do not have this type of problem and tend more to spread the alumina over the anode cover. There is more chance that the independent feeder design will ensure that alumina gets to the bath compared to the integrated feeder design.
Welch et al., (28,29) and Maeda et al. (30) deduced the effects on slug versus slow feed dissolution in laboratory cells. Roach (31) summarised the Welch/Kuschel results. All concluded that slug dosing gives quicker dissolution than slow feeding. No comparison work has been published on the effect of alternative feed designs on operating pots to verify this theory or to quantify any effects relating to anode effects, sludge formation, or current efficiency.
The crustbreaker of independent feeders needs to be cycled only when a hole is required. Hence, breaking the hole every 2-3 feeds is possible. This reduces plunger wear and reduces air usage. This is not possible for integrated feeders where the crustbreak cylinder operates the dosing unit. For integrated feeders, changes in plunger speed affect the dosing unit (refer Chapter 2). CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 28
An attraction of the independent feeder is the ability to reduce the size of the air cylinder for dosing, but still use a large cylinder for crustbreaking. Thus, the cylinder for feeding is a 50-80mm bore with 50-80mm stroke. Compared to the 125-200mm bore x 400-600mm stroke crustbreaker cylinder; this reduces the air required by a factor of 10-200. Air consumption on a potline varies, but usage of 10,000-20,000 mVl is not unusual, so the efficient operation of pneumatics is critical to an efficient feeding system and minimum use of power.
The choice of feeder design may be selected by economics. The independent feeder costs about three times that of the integrated feeder, but costs less to maintain. It is up to the company concerned to choose between capital versus maintenance costs. No published comparison tests have been conducted to determine which design is more accurate in dosing alumina or if feeder design is significant enough to affect pot performance in a measurable manner. Capital costs for a 3 OOkA pot is about $30,000 for independent feeders versus about $10,000 for integrated feeders, but feeder life is about 5-6 years versus 1-4 years for integrated feeders. Thus it is a comparison of high initial cost versus long term savings in maintenance.
The cost of maintenance is minor compared to the fact that feeder failure results in pot inefficiency, and operator exposure to falls, heat, dust and fumes. Hence, any feeder design that performs its duty with a long life is highly desirable. It would be advantageous never to change a feeder on an operating pot. Because pot life is about 5 years, people normally seek a 5 year feeder as a target. However, with the longer life being achieved by Pechiney and some other pot designs, there is a directional change to a target of a 7 year feeder. Some pneumatics suppliers are already testing cylinders of this standard.
As most plants achieving world class current efficiency and very low anode effect rates also have independent feeders, it could be assumed that the independent feeder is the best design. This conclusion may not be substantiated as there could be many other reasons that may be more significant than feeder design e.g. bath/metal height, CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 29 magnetics, alumina segregation control, shot size, operating practices, anode cover quality/variability, alumina quality, maintenance strategy, quality control.
This research has developed a better understanding of the advantages and disadvantages of the two designs and introduced trials to quantify their importance on pot performance (Chapter 10). Chapter 11 concludes that the independent feeder is the preferred design and the basis for that decision.
1.2.7 Conversion from Integrated to Independent Feeding
Considerable work on feeder design has been carried out by the author as part of this research. Two designs have the attractions of the independent feeders, are cheaper to make than independent feeders and can be retrofitted in Alcoa, Alcan and Alumax plants that already have integrated feeders (2,4).
The A2 (Appendix 6) has only one cylinder, but the feeder doses when the plunger is out of the hole, so the dose directly feeds into the hole (Figure 1-14). Improved dosing unit profile ensures more accurate dosing for variable alumina particle size. To date patents have been approved in the USA, Australia and New Zealand. One disadvantage of this feeder found in plant trials at Portland is that the feeder can jam if there is a large build-up on the plunger. This results in prevention of the dose being released and anode effects can result. There has also been a history of build-up of material on the plungers if liquid level in the pot is not kept under control.
The A3 (Appendix 8) has a small cylinder (with a hole in the centre) mounted under the crustbreaking cylinder (Figure 1-15). This hollow cylinder independently doses the pot. The retrofit costs are about half that of alternative independent feeders, yet the air savings and more accurate dosing of the latter design are achieved. To date patents have been approved in the USA, South Africa, Australia and New Zealand with approval expected in several other countries due to a very good International Examination Report by the International Co-operation Treaty examiners. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 30
FIGURE 1 -14
A2 FEEDER OPERATION
ALUMINA FLOW FROM SUPERSTRUCTURE
ALUMINA DOSE
DELIVERY OF DOSE
FEEDING
NOT FEEDING CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 31
FIGURE 1 -15
A3 FEEDER OPERATION
ALUMINA FLOW FROM SUPERSTRUCTURE
ALUMINA DOSE
DELIVERY OF DOSE
FEEDING CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 32
FIGURE 1 -16
SEQUENTIAL FEED AEDD CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 33 Clearly the A3 is considerably more complicated and consequently more expensive than the AEDD. A compromise was sought that had the accuracy and the flexibility of the A3 yet be cheaper to install. The AEDD with "Sequential Feed" (Figure 1-16) had the basic feeder shape of the AEDD but had the directional control valve (DCV) mounted on the top flange to give better control. This does not have the ability to break a hole every few feeds or the ability to deliver the dose directly into the hole, but does have the air usage savings and similar shot size accuracy to the A3 and is easy to retrofit.
1.2.8 Continuous Feeders
Another concept that has been in the development stage for a number of years is to feed smaller and smaller doses of alumina. Continuous (rather than batch) feeding in a continuous operating process appears attractive. To achieve this efficiency, one must develop a crustbreaking system that feeds often, yet does not use a lot of air. The dosing system has be unaffected by the upwards velocity of gas from the hole, and the dose must mix with the liquid bath.
Alcan, Comalco, and Pechiney have experimented with the design. Most have not proceeded with the concept, but Comalco is still working in this area and has a prototype pot in operation currently in Bell Bay, Tasmania (based on research documented in 28,29,31 by Welch et al.).
In 1990-91, the author designed a "dribble feed" chute that replaced the existing chute under the AEDD feeder to feed the pot continuously. It was possible to operate in a continuous manner for many occasions, but the design gave inconsistent results. The problems experienced were similar to those of other continuous feeders. The main problems associated with continuous feeders are:
(i) blockages at the small orifices needed for slow flow; (ii) variable flow rates if the alumina size changes; (iii) back draught from the hole in the crust blows the alumina away; and (iv) low addition rates tend to allow the alumina to float on the bath rather than mix efficiently below the surface. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 34 Though theoretically a good idea, the Portland design (and all others to date) have not proceeded past prototype testing due to these production problems.
1.2.9 Aeration Feeders
Another feeder design that was a hybrid between a continuous feeder and the independent feeder was the design used initially by Alcoa in their Tennessee smelter (32). Here an airslide was used to supply the pot directly via an air-fed manifold. This required the measured dose to flow evenly to each feeder in the pot. This design has not proved effective due to uneven flow to the individual feeders. Another aerated dosing system has now been developed for this plant via in-house development. This has proved both reliable and accurate. It is a relatively low cost independent feeder, using conventional crustbreakers to make the hole, however it uses significant amounts of compressed air.
Alusuisse developed several designs incorporating a fluidised alumina dosing unit offset from a crustbreaker unit (33,34,35). Aeration feeder designs were developed by Aluminium de Greece (36), Nippon Light Metals (37), Pechiney (35), Norsk Hydro (39), Alusuisse (40) and Alcan (41). The design that appears to be the simplest and most practical of this type of feeder is the Alcan design shown in Figure 1-17(a). Kaiser (42) used a combination of airslides and a rotary feeder to dose the alumina into the pot. These aeration feeders have not progressed past the development stage to date except for Aluminium de Greece who have retrofitted their pots with the patented design with claimed improvements of 2% current efficiency and one anode effect per day (16).
Aeration feeders suffer from the variability of alumina flow that occurs with changing particle size. Fine alumina flows easily when aerated but, when it stops, it is extremely hard to reaerate and have it flow consistently. This can be very significant if there is segregation in the supply system (which is often the case). The Moeller and Pust (43) design shows just how complicated it can be to meter alumina accurately from a fluidised bed (Figure 1-17(b)). Arnold (44) identified this as a concern with the design CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 35 of these types of feeders as part of this research project. Other concerns with continuous feeders are the build-up of scale over time, the cost of air and the difficulties of designing a reliable cost-effective air supply system to the hot parts of the pot.
FIGURE 1 - 17
AERATION FEEDERS
ALUMINA
EZZZ
AIRjh
Adapted from Patent US4,919,303 (Ref 41) (a) Alcan
LEVEL CONTROL
AIR
Adapted from Patent UK2.101,980 (Ref 43)
(b) Moeller CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 36 The author developed a feeding system based on a conventional dosing system but utilising an aeration pad to pulse the feed into the pot; the "pulse chute" (Kissane (4) and Figure 1-18). In this way, the accurate dose is achieved (due to the conventional dosing system), yet it can be metered into the pot over a longer period of time to minimise the amplitude of the concentration swings. Hence, it is a compromise between volumetric feeding and aeration feeding.
After building the prototype, it was realised that the A3 was potentially a better feeder design. Resources concentrated on the A3 and the patent has not proceeded. The main disadvantages of the pulse chute is the risk of alumina permeating through the perforated plate into the air chamber and the need to use more air to feed the pot when most smelters are trying to reduce air consumption. Appendix 7 shows the international patent application for this design.
1.2.10 Computer Control
Even with the best pot feeder available, support systems are needed to ensure the pot gets the alumina when it is required i.e. computer controlled feed logic and alumina supply systems. This has become more critical with the desire to maximise the anode area in the pot to get the greatest current possible through a pot without increasing its size. As the size of the anode increases, the bath volume reduces and the ability to achieve a constant alumina concentration becomes more difficult unless good logic is used to feed the dose at theright time.
Pechiney patented an over-feed/under-feed logic in the early 80's (23). This type of logic is now widely used to achieve the correct feed at the minimum voltage possible yet avoid the high voltage of the anode effect (which causes an upset to pot operation). Hydro Aluminium (45,(46)) and Reynolds (47) also have patents using similar logic. It is curious that the patents were awarded on almost identical concepts. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 37
FIGURE 1-18
PULSE CHUTE OPERATION
AIR IN
NOT FEEDING
'EEDING CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 38
1.2.11 Pot Alumina Supply Systems
There are basically four ways of transporting alumina to the pots.
(a) Overhead Cranes The supply of alumina to the pot has traditionally been by overhead cranes which have a hopper suspended or a hopper as an integral part of the crane. Figure l-4(a) shows white alumina dust piled up on the top of the superstructure. This is a hygiene problem to operators and tradespersons as the alumina has already been through the fume scrubbers and the dust has fluoride absorbed on the surface of the particles. This may be a contributor to asthma. Overhead crane systems have been used in many plants but some have been converted to dense phase. Plants that still have overhead crane systems include T4, T5, T8, T9, TIO, T12, T13, T14, T15, T16, T17, D2, D6, D9, D10, Dll, D19,D20,D21,andD24.
(b) Airslides Air gravity conveyors (or "airslides" as they are commonly known) have been used on several Alcoa potlines as a low maintenance design that is simple to operate. Their principle of operation is to have the alumina above a permeable fabric with air passing through to aerate the alumina. Alumina flows if the angle of the fabric is above about 7 to the horizontal. These airslides transport the alumina right up to the pot in some locations e.g. T7, T11,D12.
(c) Dense Phase The development of a dense phase system has been difficult due to build up of hard scale in the lines which reduces the conveying rate and requires the pipes to be cleared at intervals. Alesa (the materials handling division of Alusuisse) has developed a dense phase design that is used in many modern smelters as either an original installation (e.g. Portland, D3, D18, D2) or retrofitted later (e.g. T3, T6, Dl, D7, D13, D22, D23). Buhler Miag installed a system at T13 and Johannes Moeller a system at T7 but both had to be removed because of poor equipment life and build-up of scale respectively. The Alesa systems are still affected by the scale problem of an extent that varies from plant to plant. The Alesa design (48) is marketed as being cheaper than an airslide CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 39 system, easy to operate via a PLC or computer and low in maintenance cost (except for the scale...which they do not advertise).
(d) Hyper Dense Phase Late in the 1980's Pechiney developed a "hyper dense phase" for their AP30 pot (49). The system incorporates a choked airslide that is self feeding as alumina is displaced and a local low pressure region is created. It is only marketed as part of the AP30 potline package. It is similar to an airslide but works on the principle of aeration due to the reduction of pressure at an offtake. This lower pressure causes flow of alumina from head tanks. The result is a low air usage system with little or no alumina degradation.
However, the low rates do not allow removal of scale that develops in any system using reacted alumina (from the scrubbing system). Hence, they have developed a cleaning mechanism to protect the system from scale (49). This also addresses the problem of foreign objects fouling the dosing system. It allows Pechiney to explore further the continuous feeder idea, as foreign objects are a major problem with continuous feeders. The slower the rate of feed, the smaller the aperture required for flow and the more prone is the system to blockage or partial blockage which in turn stops or changes the feed rate. One can see how the supply and the feeding systems interrelate.
The hyper dense phase is very similar to a USSR patent (14) that describes fluidised supply of alumina to the sides of a pot. (Note that the USSR patent was granted several years before the Pechiney patent.) The hyper dense phase system has been installed at D4, D5, D8, D14, D16 and D25.
1.2.12 Plunger Developments
To control plunger wear, the common methods are to reduce the time the plunger is in the molten bath and/or to use stainless steel material. Generally, plants use timers in their computer logic or on local control panels near the pot to control the time the feeder is fully extended towards the bath. There is thus a fixed "dwell time" between when the crustbreaker descends to when it is retracted for most (if not all) pots. The time that the plunger is immersed in the bath is therefore dependent on the response CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 40 time of the valves and pneumatics and on the speed of the piston in the crustbreak cylinder.
To minimise the time in the bath, several designs have been developed. Alusuisse has a sensing unit that detects a voltage change across the feeder when the plunger strikes the bath (52) so the plunger is retracted quickly (Figure 1-19(a)). Pechiney uses a similar concept on a small scale crustbreaker to detect liquid level for bath level control in their AP30 pot (19,53) and illustrated in Figure l-19(b).
The problem with the Alusuisse and Pechiney designs is that they require the plunger to get wet to work. It is not necessary to do this. A good operating practice is to control the liquid level below the plunger when the plunger is fully extended. In this way, the plunger does not get wet and the plunger life is extended significantly. In Chapter 5 this philosophy is discussed in more detail.
A better device to minimise bath corrosion of plungers is end-sensing of the crustbreaker cylinder. This design is marketed by some cylinder manufacturers e.g. Atlas Copco and Mechman. Pneumatic designs are used due to the effect of the strong magneticfield associated around operating pots. Pot magnetic fields are not only strong but also variable. This can seriously affect the performance of magnetic sensors. However, pneumatic sensors are seldom used due to their high cost and the large number of feeders per potline. A pneumatic design was investigated as part of this study (Chapter 10).
In the absence of an end-stroke sensing, the next option is to have a plunger that lasts for long periods. The only published work is that of Kaiser in 1984 (54). Strips of different analysis cast materials were welded to a cast iron plunger in a test pot to gauge relative wear rates. The conclusions made were that 20% chrome was required to give reasonable wear rates and that the cost was economic on a cost per tip-day basis. (In Chapter 5 this paper was examined and a much better relationship was extracted from the data.) The conclusions of the Kaiser paper have been adopted by the industry and most plants have changed from cast iron to high chrome plungers. However, plants continue to experiment with different plunger materials. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 41
FIGURE 1 -19
PLUNGER END STROKE SENSING w
Adapted from Patent AUB-24373/84 (Ref 52) (a) Alusuisse
Adapted from Patent AU14784/88 (Ref 53)
(b) Pechiney CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 42 The mechanism of plunger wear has been under discussion since pots were invented. Some believe the cause is erosion and others believe it is corrosion. This study systematically identified (by theory and plant trials) the causes of plunger wear and examined 19 materials and different shaped plungers (Chapter 5). The economics of different plunger materials and designs are discussed in Chapter 9.
1.2.13 Dissolution of Alumina
A key factor in the development of a better pot feed system is to understand the properties of the alumina...both its physical flow properties, and the ability to dissolve efficiently. Alcoa, Alusuisse, Mitsubishi, Comalco and Hydro Aluminium have either researched in-house or engaged expertise from universities or research establishments to investigate these areas. Pechiney has not published their "in-house" research.
Most published literature on alumina dissolution has been generated by Welch (University of Auckland) (28,29,55,56,57,60,61,62,63), Bagshaw (formerly Alcoa) (28,29), Kuschel (Pasminco Research Centre, Newcastle) (56,57) and Taylor (Comalco) (60,62). Several of these papers refer to similar results and conclusions. For four years, Alcoa sponsored this research (summarised by Roach (31)) then Comalco engaged Welch to continue research into dissolution and to develop a continuous feeder that was operating in Tiwai Point (New Zealand) before changing to Bell Bay (Tasmania) in 1992.
Most alumina used in pots is that which has passed through scrubbers that treat the exit gases from the pots...called "reacted" alumina. This is ingrained with cryolite, SO2 and fluoride. It is sticky and does not flow like fresh alumina. Many papers published on the reactivity of alumina by Welsh et al. (28,29,55,56,57,60,61,62,63) have used fresh alumina. This may give false impressions of what actually happens in operating cells. This is a major fault with their published studies, and casts doubts on their results.
Jain et al. (55), Welch and Kuschel (56) and Barrillon (58) discussed dissolution studies and the equipment used for the experiments. Taylor et al. (60) and Thomstad et al. CHAPTER l INTRODUCTION AND LITERATURE SURVEY PAGE 43
(61) explained the effect of sludge and showed how difficult it is to dissolve once formed. Modelling has been used to simulate the effect of dissolution in cells and cell geometry by Liu et al. (62) and Haverkamp et al. (63).
The dissolution work of the University of Auckland tends to support results of Maeda et al. (30) that fast addition of alumina gives fast dissolution. Also, the predominance of fines does not enhance dissolution (although one would expect it should give a faster dissolution). This is due to non-wetting of the dust. The resulting 'raft' of alumina floats on the bath surface and does not break up into individual particles easily. Agitation of the liquid is necessary to physically break up the rafts.
Roach (31) identified the order of priority of factors that affect alumina dissolution rate (most important first):
1. addition systems; 2. bath agitation; 3. superheat; 4. alumina concentration; and 5. alumina properties.
Alcoa's Western Australian Refineries considered it was up to the smelters to get control of the feeding system before more work was needed on identifying alumina chemical factors. This led Alcoa to cease engaging outside consultants into dissolution, although Alcoa's very substantial Research and Development Department at Kwinana (Western Australia) continues to develop alumina quality. Hence, from all the above work, it can be concluded that it is best to have good feed control and that the pot
feeder clearly has a large impact on sludge (muck) control.
The type of feeder dose will have an effect on dissolution rate. Purdie (103) showed in
experiments with operating pots that about half of the alumina feed in pots fed by bar breakers comes from the bottom of the pot. Other pots fed by independent point
feeders only "backfed" 10-20%. What is not quoted in literature are comparisons of different feeders in the same pot. This is the only way to measure the significance of
integrated versus independent feeder designs. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 44
1.2.14 Alumina Dose Design and Accuracy
The Pechiney API8 independent feeder is claimed to achieve a feed shot size "accuracy within 5% of the designfigure" (20). For a mass of 1kg, this relates to +/-50g. This is the only published data source that specifies the accuracies of pot feeders. Discussions with Holmsi and Leonard (101) verified this to be typical of Pechiney pots. It is reasonable to assume that this is the accuracy of other independent feeders as they are modelled on the Pechiney feeder. The Boyne integrated feeder accuracy is +/-100g (102) for a 1.8kg shot and the AEDD accuracy is about +/-40-250g. Chapter 2 shows results far superior to these for the sequential feeder and the A3 feeder viz. +/-14g and +/-12g respectively. The accuracies of the point feeders compares very well with the Soderberg and bar break designs where 30-70kg shots were used.
To address the effect of segregation and alumina size variation, some dosing systems are designed with approximately 45 angled surfaces at top and bottom. This is seen in designs of a USSR patent (24), the Pechiney API 8 feeder (22,23) and Figure 1-12.
Arnold (44) and Richards et al. (64) conducted tests on flow and angle of repose of reacted alumina with relevance to feeder dosing units. Richards showed that the angle of alumina at the top of the dosing unit varied depending on its size and fluidisation; fine material had a stationary angle close to horizontal and coarse material was at a 45 angle. Arnold (44) recommended that the exit angle should be as steep as possible (at 60 or steeper) to ensure all alumina exits the dosing unit. Kissane (76) combined these two observations and developed a cheap ($40) spool insert with 60 angles at top and bottom that could be retrofitted into the AEDD integrated feeder (Figure 1-9 and 1-11), and incorporated this design feature in the A2 (Figure 1-14) and A3 feeders (Figure 1- 15). Thus, alumina size and degree offluidisation would have a minimum effect on dosing accuracy. Data to support this is presented in Chapter 2.
The flow properties of alumina have been more tightly defined recently by Pechiney both from a chemical and physical viewpoint (65). This has become necessary as feeder development has highlighted how important the particle size of alumina is to pot feeding. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 45 Reverdy, a Director of Pechiney, states that direct entry into the hole is a benefit of independent feeders (66). It is interesting that for some Pechiney plants, the crustbreaker only operates every 2-3 feeds e.g. D10. This prevents direct entry to the bath and suggests that direct feed may not be a major factor in pot control. To identify the effect of direct feed on dissolution, this research developed chutes under the feeder that allow direct feed into the hole. Some of these can be seen in Figure 1-14 and 1-15. Section 2.8 discusses the effect of chute design on the pot.
Patents from the USSR have explored different concepts in feeder design from integrated to independent feeders. Chevonin et al. (67) have developed a dosing hopper with an internal wire brush to clean out alumina. This is a costly solution to segregation when one considers there are a couple of hundred feeders in a modern smelter. Mirkin et al. (68) have a patent for a double valve dosing unit controlled by sieves sequentially releasing the dose into the pot. This design appears to reduce internal by-pass in the feeder, but may still have poor accuracy due to leakage around the valves. Costs are expected to be high.
1.2.15 Pneumatic Air Cylinders
Pot crustbreaker and feeding units only use pneumatic cylinders as there is a potential risk that hydraulic oil may createfires when exposed to the elevated temperatures of a pot; up to 1,100°C is possible. Under the varying conditions of pot temperature, the desire to last over 1 million cycles (about 5 years) historically has proven to be a difficult assignment. The causes of failure of cylinder and assembly components are similar in most plants, but the extent differs due to component design, to downstream failure from other parts, to pot conditions and to repair/installation errors.
Traditionally, smelters have sought a cylinder with the longest life, rejecting all that do not achieve the desired results. For some plants, this has resulted in many years of trials of different cylinder designs. This study has shown that many factors affect cylinder life, not the least of which are the pot itself and the feeder or crustbreaker design (Chapter 7). CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 46 Figure 1-20 illustrates the main components of a typical pneumatic cylinder. Each cylinder brand has different configurations of seals (that prevent air leakage past the piston rod or barrel), bushes (which prevent the shaft from moving transverse to the centre line), cushioning (to slow the piston at the end of its travel) and seal/bush orientation. The cylinder shown is an Atlas Copco cylinder that has one of the best performance history in pot feeder operation.
FIGURE 1 - 20
PNEUMATIC CYLINDER PARTS
Adapted from Atlas Copco, Series 40/41M brochure, 91270041-92, 2nd edition, Pp,3 CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 47
No papers or patents could be sourced on comparison of different cylinders nor any other system components associated with the cylinders e.g. exhaust mufflers, DCVs, flow control valves. The performance of different components can only be found by discussion with suppliers to plants or with representatives of the plants themselves. Because of the long life of feeders, transferring of engineers and maintenance personnel, and the lack of failure data, there are little valid data from most plants.
In order to establish optimum designs and suppliers, this study contacted many smelters (Table 1-VI). Plants with integrated feeders generally had poorer performance to independent feeder plants; 1-2 year life versus 5-6 years. Independent feeder crustbreaker units generally operate at slower speeds to integrated feeders, so this reduces frictional wear on seals and bushes. Independent feeders that break a hole every 2-3 feeders have an added advantage over most plants by having less cycles per year. Examination of the plant data shows that there are extreme inconsistencies in the performance of the one type of cylinder in two different plants (Chapter 7).
For example, of the 16 integrated feeder plants contacted, T17 had the longest integrated feeder cylinder life at just under 5 years. This plant uses Scheffer 100mm O.D. cylinders, but the key to the good performance was retrofitted Dover Teflon seals. Identical cylinders (without Dover seals) at T7 and T15 only achieved 1-2 year life. Use of 20 similar seals at Portland showed life under 1 year. The longest life cylinder of the 8 independent feeder plants contacted was from over four hundred Atlas Copco cylinders at Karmoy (93) which theoretically had a life of 94 years! (Table 1-VTI). Yet a 150mm O.D. cylinder on an integrated AEDD design at T2 only had a 2 year life.
Given that most plants are not immediately in a position to change their large number of cylinders (up to 2,000), there are still opportunities to improve cylinder life. Plants T7 and T16 addressed the source of feeder failures and achieved a 65% and 90% drop in failures respectively, and an 80-90% drop in costs in two years, yet still use basically the same cylinder and plunger designs as they did originally. As will be seen Chapter 10, Portland feeder and cylinder life improved by 90% as a result of this investigation. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 48
TABLE 1 - VI
SMELTERS CONTACTED IN THIS RESEARCH
Capacity Integrated/ Location Smelters/Research Centres Visited AEDD (t/year) Independent
Australasia Boyne Smelters Limited (Comalco) 230 integrated Bell Bay (Comalco barbreak) — barbreak * Tomago (Pechiney design) 380 independent Kurri Kurri (Alcan) 153 integrated * * NZAS (Comalco) 269 barbreak * Pt Henry (Alcoa) 180 integrated * * Portland Aluminium (Alcoa) 320 integrated * *
South America Alumar, Brazil (Alcoa) 350 integrated * * Aluar, Argentina 175 independent
U.S.A. Wenatchee (Alcoa) 220 integrated * * Badin (Alcoa) 115 integrated * * Massena (Alcoa) 127 integrated * * Rockdale (Alcoa) 310 integrated * * Warrick (Alcoa) 270 integrated * * Tennessee (Alcoa) 200 independent * Alumax, Mt Holly (Alcoa design) 181 integrated * * Kaiser Research Centre, California — — Mead (Reynolds) 200 integrated Vanalco (Alcoa design) 115 integrated * * * Europe, UK Mosjoen, Norway (Alcoa design) 120 integrated * Granges, Sweden (Reynolds) 100 independent * Karmoy, Norway (Hydro) 220 independent * Hydro Research Centre, (Norway) — — Lynemonth, England (Alcan) 130 integrated * * Anglesea, England (Kaiser) 117 barbreak ISAL, Iceland (Alusuisse) 92 independent Ardal, Norway (Alusuisse) 188 independent * Canada Alcan Head Office, Montreal — Le Terriere (Alcan) 204 integrated * Baie Comeau (Reynolds) 400 independent * TOTAL 5,366 2,995 (26% of western world capacity) Note: Installed capacity from Pawlek (94);updated 1994. Suppliers Contacted Parker-Hannifin, Sydney cylinders Terry Fluid Controls, Melbourne cylinders SMC, Sydney cylinders Atlas Copco, Sweden cylinders CPOAC, France cylinders Norton, U.S.A. seals Dover, U.S.A. seals CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 49
TABLE 1 - VTJ KARMOY CYLINDER REPAIR HISTORY
Hydro Aluminium, Karmoy Fabrikker, Elektrolyse Prebake, Karmoy. 21.10.92
To Jim Kissane
According to your questions yesterday about the Atlas Copco Monsun Tison cylinders type; C 40-200-50-500 F 00 installed in 1987, we can give the following data;
Number of Atlas Copco breaker cylinders failures
in 1989 2 numbers in 1990 11 " in 1991 12 " in 1992 8 "
For your other questions I am not allowed to answer because of the technology agreement.
Kind regards Magne Stueland,
Editorial note: Karmoy have 444 feeders. The average cylinder life for a turnover of 8.2 per year is calculated at 94 years versus 2-5 yearsfor mos t plants. Note that the quoted data only relates to cylinders and not to feeder or plunger life. Evenfor cylinder life, these results are outstanding. The "other questions" (mentioned in the last sentence of the letter above) were relating to plunger and feeder information. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 50
1.2.16 Pneumatic Systems
The invention of point feeders reduced the shot size from 30-70kg to l-2kg. As pot size increased, the number of feeders per pot increased from one per pot to a maximum of five per pot (Figure 1-21). To distribute this alumina across the pot, different air delivery systems (commonly called "pneumatics" in this text) have been used.
In Alcoa/Alcan/Alumax plants, the air supply to multiple feeders/pot have been by one directional control valve (DCV). The independent feeder systems require the crustbreaker and dosing units to be separately controlled, so two DCVs are required per feeder. In order to reduce the air usage, it was desirable to mount the DCVs as close to the feeder as possible, so the valves were generally positioned on the pot superstructure. (It will be shown in Chapter 2, 4 and 5 that this also improves shot size accuracy, pressure at the crust and plunger wear.)
Air usage for pot feeders is quite significant e.g. about 30% of Portland's usage is from pot feeders. For a smelter, the cost of air compression may not just be the cost of power and maintenance of the compressor equipment. There can be a very expensive "cost penalty" as generally there is a maximum power limit on the plant. It is highly desirable to maximise the availability of power to the pots. The more pots that can operate, the more profit can be made from metal sales.
Compressors and other uses of "ancillary power" have a cost penalty which is several times the normal calculated air compression cost. The degree of "penalty" will vary depending on whether one can sell the metal or not. If the metal cannot be sold due to a oversupply (as has been the case for most plants in 1994/1995), then the cost of air is the prices quoted above. If metal can be sold, the "penalty" can be several times these values.
Not only is the cost of air generation a factor, but also the available quantity of air and the net pressure that is available. Often, overuse of air leads to low supply pressure with subsequent effect on crustbreaker plunger pressure. Dunstan (69) proposed changing feeders based on air consumption. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 51
FIGURE 1 - 21
PORTLAND POT FEEDER SYSTEM
WIDE OR TAP AISLE
RESTRICTOR (12mm)
NOTES: (1) In closed position, ball valve NARROW has a 12 mm hole. AISLE (2) Line A supplies air to force plunger up. (3) Line B supplies air to force plunger down. RESTRICTORS (12mm)
IR SERVICE UNIT -BALL VALVE (BLEEDS TO ATMOSPHERE) -FILTER -LUBRICATOR GATE VALVE SUPPLY AIR CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 52 It is considered that this is not an effective manner of cost reduction and that the particular problems at the plant concerned should be approached in a different manner rather than accept that excessive air leaks are inevitable. As no other study of maintenance and cost reduction has been published, Chapter 8 examines the causes of crustbreaker/feeder failures and systematically identifies opportunities for plants to increase equipment life. Chapter 9 examines the economics of different maintenance strategies and feeder designs.
To minimise the high air costs of feeding systems (especially the crustbreaker) some smelters utilise pressure regulators or commercially available latching systems to hold the air cylinders in the up position. Swiss Aluminium has patented a dual pressure setting system for the cylinders to keep low pressure on the feeder when in the up position (70). By reducing this pressure, there are cost savings due to less air usage. Ross Operating Valve Company market a similar device that is claimed to reduce air usage by 50% by a similar method, but there are few units operating in plants at present. Terry Fluid Controls and Comalco jointly developed and have applied for a patent for a pressure regulation system that also reduces the pressure holding the cylinder retracted to 100-200kPa. The claimed benefits are a reduction of up to 50% in air leakage rate and faster downstroke time (98). The units will be on trial at several smelters (including Portland) by end of 1995.
All four of these designs accept that air leaks are inevitable. However, these designs may be uneconomic if seals do not leak or do so to a minor extent, as each feeder needs a unit worth several hundred dollars. It appears to be more efficient to address the primary cause of leakage rather than work on the effect.
1.2.17 Other Feeder Developments
Alcoa has worked on a gravimetric feeder that avoids the bulk density change resulting from a change of alumina size, but this has not led to plant scale installation due to high costs (71). As will be seen in Chapter 2, this research discovered that the variation of bulk density had only a small effect on shot size accuracy provided size changes are not CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 53
extreme. Also, this research identified that accuracy from shot-to-shot for volumetric feeders on operating pots can be as low as 0.5% of shot size, which is sufficiently accurate not to consider costly alternatives and is far better than the 5% accuracy quoted by Pechiney (20).
The gases given off by anode effects are not desirable to people who work in the potroom or operating areas (1). Greater awareness of the health and hygiene of operators and tradespersons in potrooms has driven managers to address exposure of people to heat, dust and fumes which are faced during change-over of feeders, unblocking feeder holes, and anode effects (72,73). Asthma is of concern in most smelters as this may be aggravated by feeder monitoring, changeout or servicing where people are potentially exposed to dust or fumes (74).
In recent years, the reduction of anode effects has become a high priority for aluminium companies across the world as anode effects may lead to global warming. However, recent research has questioned the relative magnitude of the effect of aluminium smelting on global warming. Studies between the environmental protection agencies and aluminium producers world-wide are assessing the impact of aluminium smelting on global warming (95). Clearly this is an area that is the subject of much research.
Irrespective of the relative significance of the effect of anode effects on global warming, there are many reasons why it is useful to optimise pot alumina feeders...cost control, pot stability, health and safety of personnel. The financial benefits alone are substantial. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 54 1.3 DEVELOPMENT OF THIS RESEARCH
1.3.1 Research Objectives
The research objectives of this project were to:
(i) optimise existing feeder designs (especially the AEDD feeder), (a) reduce failures (b) reduce anode effects (c) reduce costs
(ii) compare integrated and independent feeder/crustbreaker units (called "feeders" in this text) and develop a state-of-the-art feeder which has the following features: (a) life of over 5 years, (b) cost effective in both capital and maintenance costs, and (c) can be retrofitted in integrated pots easily.
1.3.2 Optimizing Existing Feeder Designs
In 1990, the author commenced optimising the AEDD feeder at Portland, but with a view that this research would have a wider application across Alcoa and other plants in the industry. Contact with other Alcoa smelters showed there was a wide range of crustbreaker/feeder life being achieved across all 9 plants, despite having almost identical designs (Table 1-IV). The distribution of causes of failure varied significantly from plant to plant (Table 1-V). Thus, plant to plant differences seem to be significant even with the same basic design.
Early in the investigation, it became apparent that the conditions at Portland were much worse than many (if not all) Alcoa and non-Alcoa plants e.g. hotter, multiple feeders per air valve, high speeds, lubrication breakdown, inferior cylinders, eccentric mountings, design faults, deep penetration in the bath. It was considered by several pneumatic cylinder manufacturers (Atlas Copco, Terry Fluid Controls, Parker Hannifin) CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 55 that the conditions at Portland were the worst seen in their experience...especially the temperature exposure. Thus, optimising Portland feeders was indeed a challenge.
An extensive literature search on feeder technology (mentioned previously in this chapter) showed that feeder developments are often kept secret within many organisations. Often published literature contains information that is either already widely known or is published as a marketing strategy e.g. Pechiney. For this reason, except for dissolution studies conducted by a handful of university researchers, the majority of published work was in patents. It was apparent that to get improvements in feeder technology, one had to do the development oneself, using personal contact with other plants wherever possible to flesh out potential problems.
It has been the custom in the aluminium industry that different companies do not assist each other, very much due to the competitive nature of the industry. From 1990 to 1994, the author attempted to break this down by contacting by phone and faxcimile many non-Alcoa plants who also were frustrated that similar feeder problems were experienced. Representatives from these plants saw that there was an opportunity for all parties to benefit jointly from cross pollination of ideas. Often confidentiality concerns restricted data or information being shared, but comments on cylinder design and feeder performance were usually freely given.
The author visited 18 smelters to inspect and discuss feeders (9 integrated, 7 independent, 2 barbreak) The plants and suppliers visited are listed in Table 1-VIData and input from up to 27 smelters are presented in this report; 17 integrated, 24 independent, 1 Soderberg, and 2 guillotine. The plant name is generally coded to protect confidentiality unless the data is on public record or presented as a formal correspondence. A "T" code is used for plants who use integrated feeders, and "D" is used for those who use independent feeders (as mentioned in the Note to the Reader on page IV). CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 56 The author approached optimising existing feeder designs by extensive use of statistics and economic analyses. By use of chemical engineering as well as mechanical engineering principles, pot operation and feeder maintenance aspects were investigated. Close co-operation with operators and tradespersons was used as a means of implementing designs that were applicable to all plants, but not designs that only had attractions on paper...they also worked in the field. Designs that assisted the safety of operator and tradespersons were developed in association with improvements to costs and feeder life.
There are common items in integrated and independent feeders that this research identified....cylinder design, plungers, dosing units, air supply hardware. The investigations undertaken resulted in an understanding of "crustbreaker/feeder technology" in general, rather than just the technology peculiar to Alcoa AEDD feeders. In this document, suggestions are made on optimising the design of all existing feeders as each design and plant location has the potential to improve performance.
1.3.3 New Feeder Designs
Contact with plants using independent feeders identified usually much better feeder life and lower maintenance costs than plants using integrated feeders. However, the capital cost is several times higher and the design cannot be retrofitted in pots currently using integrated feeders.
Hence, this research developed designs that could be retrofitted into plants that currently use integrated feeders. The designs were to be cost efficient alternatives that had the attractions of the independent feeders, but not the same high capital cost. As noted previously, the A2 design has been successfully patented in Australia, New Zealand and the USA (Reference 2 and Appendix 6) and the A3 in Australia, New Zealand, South Africa and the USA (Reference 3 and Appendix 8). The A3 has the potential to achieve independent feeder performance (or better) and still be able to be retrofitted into existing AEDD feeder pots with minimum installation costs. Figure 1- 22 illustrates the feeder designs invented as part of this research along with the AEDD feeder. CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 57
FIGURE 1 - 22
PORTLAND FEEDER DESIGNS
AEDD A2 PULSE CHUTE SEQUENTIAL A3 FEED CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 58 1.3.4 Across Alcoa Optimisation
It is useful to explain the process on how Alcoa plants responded to, and assisted in, this research.
Each of the nine Alcoa smelters world-wide has a team of people investigating improvements in key activities that are considered the major items to make Alcoa world class in all aspects of aluminium production. These teams are called "In Control and Capable" (ICC) Teams and the leaders from each make up the ICC Network.
The Feed Alumina Network addresses the control of anode effects and pot control. Massena (USA) is the lead plant for this Network and was responsible for benchmarking the best in the world. In 1991, all ICC teams submitted their key findings to senior international Alcoa management. One of the Massena recommendations was that Alcoa should "support feeder technology development in Portland". Later, Alcoa listed "Portland Feeder Technology Development" as one of the 39 enablers for Alcoa to become world class; only two items were listed for Australia.
In April, 1992 at Wenatchee (USA), the Alcoa Feed Alumina Network hosted a presentation by the author on feeder technology to discuss opportunities from which each plant could benefit. A list of ten general items was raised to be considered by all plants; commonly called the "Kissane Feeder Recommendations" (Table l-VlII).
Appendix 4 (reference 99) is the instruction from Bob Seymour (Alcoa Quality Manager, Primary Metals) to the Alcoa Smelting Managers to implement the feeder recommendations in all plants and why they are important. These recommendations became one of the seven Critical Activities for Alcoa smelters for the next two years and implementation became a high priority item in all nine international Alcoa plants. CHAPTER l INTRODUCTION AND LITERATURE SURVEY PAGE 59
TABLE 1 - VIU ALCOA FEEDER RECOMMENDATIONS
1 TRACKING SYSTEM 2 SUPPLIER QUALITY ASSURANCE 3 SURVEY DOWNSTROKE TIME 4 SURVEY BLOCKED FEEDER HOLES 5 REDUCE DWELL TIME 6 DUAL DWELL TIME 7 PERFORMANCE CRITERIA 8 SPOOL INSERTS 9 PLUNGER COMPOSITION/SHAPE 10 PISTON LEAKAGE SURVEY
(PLUS INDIVIDUAL PLANT RECOMMENDATION LISTS)
After two years, the recommendations were removed from the critical activities list due to the success of feeder research (as outlined in Appendix 5 (reference 100) by Dick Taylor (Director of Technology, Alcoa Primary Metals)). Teams are currently operating at most locations to optimise pot feeders and they are continuing to achieve very worthwhile results.
1.3.5 Document Layout
To become knowledgeable on the state-of-the-art, a literature survey was conducted as well as contact with other smelters and equipment suppliers (Chapter 1). This research consideredfirst of all the desire to measure an accurate dose (Chapter 2) and deliver it into the bath (Chapters 3). A critical factor in the operation of feeders is the air supply system. Chapter 4 describes tests conducted on the pressure profiles of cylinder inlet and outlet ports to establish what affects the speed and pressure of crustbreaking units.
As plunger failures are one of the predominant causes of crustbreaker unit removal, plunger wear and build-up on the plunger were discussed in depth (Chapter 5 and 6). CHAPTER 1 INTRODUCTION AND LITERATURE SURVEY PAGE 60 No matter how accurate the feeder is, a requirement of maintenance personnel is to contain costs and operating personnel do not want to be troubled by frequent changeouts. One of the most frequent causes of failure is the pneumatic cylinder, so Chapter 7 discusses extensive trials on cylinder operating variables and cylinder lubrication.
The causes of failure of feeders is discussed in Chapter 8 and recommended solutions are raised to extend feeder life. Chapter 9 establishes that long life can be complemented by cost efficient designs and maintenance strategies to minimise feeder costs. Note that the two may not go hand in hand on a short term basis, but, on a long term basis, long life generally results in reduced costs.
An integral part of this research was not just to make existing feeders better. Several designs were developed and the good results indicate that these designs of low cost feeders are capable of achieving the targets of long life and enhance the operation of the pots.
Finally, Chapter 11 collates the results of the feeder research both in Australia and overseas, and indicates where the technology can improve in the future. CHAPTER 2 SHOT SIZE ANALYSIS PAGE 61
CHAPTER 2
SHOT SIZE ANALYSIS
2.1 SUMMARY
About 200 tests involving over 30,000 measurements were conducted on the shot size of several pot feeder designs viz. the existing Portland AEDD, the A2 and A3, plus two feeders from plants T6 and Tl 1. These tests sought to establish the factors that affected shot size mass and variability. Parameters considered were dwell time, stroke time, alumina size, feeder-to-feeder differences for the same design, cylinder brand, spring type and cylinder cushioning quality.
Portland AEDD feeders were significantly affected by many parameters with the most important being feeder to feeder changeout differences, alumina size, type of cylinder, and dwell time. The overall result was a potential variation in shot size of up to 37% and a standard deviation of up to 14% shot-to-shot. There was an improvement of 56% in shot-to-shot standard deviation and 36% in feeder-to-feeder standard deviation when a 60 spool insert was installed for a changeover cost of $40.
The standard T6 and Til feeders were better than the Portland AEDD but were affected by the same parameters. The A2 was more accurate with a lower standard deviation than all the integrated feeders, and was less affected by the parameters considered. The A3 independent feeder had a standard deviation of about 0.6% of shot size for shot-to- shot accuracy, making it by far the most accurate feeder tested and more accurate than Pechiney independent feeders.
The improved shot size accuracy due to installation of spool inserts on Portland AEDD feeders was a major enabler to a significant reduction in anode effects. It was established that up to 32 parameters can affect shot size accuracy for integrated feeders and 20 parameters can affect independent feeders. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE62
2.2 INTRODUCTION
A study of shot size mass from pot feeders was initiated in December 1990 to determine what factors affected shot size mass and accuracy. Several feeders were tested:
- existing Portland AEDD - Portland A2 - Portland A3 - T6 integrated feeder -Til integrated feeder
Also, individual and multi-feeder operation per DCV were tested to see the effect of pneumatics on shot size accuracy.
The A3 is an independent feeder that was expected to give similar results in shot size accuracy to other independent feeders, as the effective piston area and stroke distance are similar to those of most independent feeders. This should give a good comparison of integrated and independent feeder dosing accuracy.
An insert was invented by the author to install inside the dosing spool on the Portland feeder following examination of results of alumina flow properties by Richards et al. (64) and Arnold (44). This design could be used in virtually all Alcoa, Alumax and Alcan integrated feeders and was tested in Portland, T6 and Til feeders.
The Portland feeders were tested using 125mm O.D. x 508mm stroke Parker, Terry Fluids, and Atlas Copco pneumatic cylinders. The T6 and Til feeders used 100mm O.D. x 356mm stroke cylinders of different manufacturers...Terry and Lindberg respectively.
Different stroke times were chosen for each feeder to be consistent with field testing on operating pots. CHAPTER 2 SHOTS SIZE ANALYSIS pAGE 53
An off-site testrig was constructed based on the Portland superstructure dimensions (Figure 2-1). Tests were conducted using normal and fine sizes for both pure and reacted alumina. Alumina used in most smelters is not fresh alumina direct from refineries. The alumina is generally passed through reactors that scrub pot gases before being transported to the pots. This affects its flow properties and dissolution rates.
FIGURE 2 -1 OFF-SITE SHOT SIZE TEST RIG
Most papers published on the dissolution of alumina in cells have used unreacted alumina (refer Section 1.2.13). This may give a false impression of what actually occurs in pots as unreacted alumina has better flowability than reacted alumina. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 64
Reacted alumina has a sticky nature due to absorption of exhaust gases from the pots. It is indeed possible that dosing unit performance could be quite different for unreacted and reacted alumina. Tests of feeder shot size accuracy should use reacted alumina as this is what is used in the pots. Hence, reacted alumina was used in this study except for comparison tests at the start of the investigation. Alumina was collected from the superstructure of off-line pots (with the exception of these early runs).
Each shot of alumina was caught in a plastic bag around the outlet chute with minimal dust lost. Shot size masses were graphed on control charts as they were measured (Appendix 1) and summary data (mean mass and standard deviation as a percentage of shot size) were calculated (Appendix 2).
Appendix 1 illustrates raw data from one typical run. This illustrates that the shot size was in statistical control as the range and mean values were within the control limits. The bottom curve is the height of alumina in the hopper around the feeder. This shows no correlation between the shot size and height. Thus, there was always enough alumina to supply the feeder for these tests.
During each run of 25 shots, samples of alumina were taken by catching several shots. These were placed in sealed containers and analyzed by EML Laboratories for sieve sizing, Malvern micro sizing and "funnel flow time". Some of the early runs were also analyzed for loose and packed bulk density (Appendix 3).
Alcoa has established a linear relationship between microfines content (-20 micron) and the time taken for afixed quantity of alumina to flow through a standard shaped funnel. The finer the alumina, the longer the "funnel flow time" (FFT). This is a simple repeatable test to compare fines content whereas expensive microsizing equipment may be affected by machine-to-machine variation. When one considers that a point feeder is largely affected by flowability of alumina, the FFT is a very useful parameter to compare alumina types. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 65
The early tests identified several factors affecting shot size that had not previously been understood as being important viz. cylinder speed, spring tension, cylinder design, cushioning quality and dwell time. These factors were examined in later tests. Tests on operating and off-line pots were conducted to verify the off-site tests and to examine the effect of "multi-feeder per DCV" and "one feeder per DCV" operation on shot size accuracy.
Some 30,000 shot sizes were measured in about 200 runs. A full detailed analysis of the results of this investigation is contained in a report by the author (76).
2.3 PROCEDURE
2.3.1 Dwell Time
"Dwell time" is the time from when the computer energises the DCV to when it deenergises the DCV. It includes the stroke time of the cylinder plus the time the piston/piston rod is stationary in the extended position. For integrated feeders, the dwell time is selected to achieve two objectives:
(i) all the feed shot is discharged from the dosing unit; and
(ii) sufficient pressure is achieved for the plunger to break the crust.
For independent feeders, a different dwell time is used for the feed and crustbreak cylinders, however the objectives are the same as for integrated feeders but split into function (a) for the feed cylinder and (b) for the crustbreak cylinder.
Dwell times from 0.5s to 5.0s were used in a random manner for each test to avoid any systematic errors. Repeatability was established by measuring the mass of 25 consecutive shots for the same dwell time conditions. There was no correlation between shot size and order of testing in any testing done off-site or on-site. Figure 2-2 illustrates the relationship between shot size mass, shot size standard deviation and dwell time for the existing Portland AEDD feeder. This type of figure has been used throughout this chapter to illustrate the performance of different designs for varying dwell time. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 66
FIGURE 2 - 2
PORTLAND AEDD SHOT SIZE GRAPHS PARKER CYLINDER.lOs down,1.1s upstroke )
SHOT SIZE (g) 2000
#1 AEDD FINE #1 AEDD NORMAL #1 AEDD UNREACTED
WRUV «»(•»*•» • • a a •
#2 AEDD FiNE #2 AEDD NORMAL
200 -L 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 DWELL TIME (s) STD DEV (%) 12 10 - 8 6
4 ""Htra-iio* 2
RUNS 1,7,13,19,22
Note: 1.7/1.9 refers to 1.7s downtroke and 1.9s upstroke time. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 67
2.3.2 Dosing Unit Shape
It was found from graphs such as Figure 2-2 that the feed shot varied with dwell time. This suggested that the shot size increases the longer the dosing unit is stationary. It appears that the alumina does not flow out smoothly (by mass-flow) from the dosing unit. This is a phenomenon that could affect both integrated and independent feeders, but is less critical for independent feeders which may use long feeder unit dwell times as the dwell time does not affect the time the plunger is potentially in the bath.
This variation in shot size was alarming in its magnitude but not particularly unexpected based on test work conducted on flowability of reacted ore. Arnold (44) suggested that at least 60-70° angles (not 45°) should be used on the dosing spool for mass-flow. The original outlet angle of all Alcoa, Alcan and Alumax feeders and most Pechiney and independent feeders is 45°. The Pechiney and other independent feeders have a 45° angle at the top as well (as illustrated in Figure 1-11).
Richards (64) also suggested that there was variability due to the angle of repose of various alumina sizes. Coarse alumina has a 55 angle to the horizontal and fine alumina has almost a 0° angle to horizontal (especially if aerated). These findings were
also confirmed by Arnold (44).
Thus, both inlet and outlet angles are important to shot size variability. The angles of independent feeders should be increased from 45°, and Alcoa/Alcan/Alumax feeders should have an angle at the top as well as at the bottom for better control over shot size
variability.
On the basis of this information, an insert with 60° angles at top and bottom of the dosing unit was designed to improve variability of shot size. This insert can be installed inside the existing spool of the Alcoa/Alcan/Alumax integrated feeders, as most have
identical dimensions. Thefinal insert designed at Portland is seen in Figure 2-3. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 68
FIGURE 2 - 3
SPOOL INSERT ASSEMBLY
(a) Disassembled
BASE
(b) Assembled (c) Details
Note that it is important to install the insert the right way up. If it is installed upside
down it restricts the outlet port of the dosing unit and can cause very long discharge
times and a further 20% drop in shot size (Figure 1-9 and 2-3). CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 69 2.3.3 Stroke Time
The stroke time is the time that the piston or plunger takes to travel from fully retracted to fully extended (or vice versa). For simplicity, the downstroke and upstroke times are reported together; such as 1.0/1.2s for downstroke/upstroke times.
In 1991, investigations commenced in reducing the stroke speed to reduce fatigue failure of piston rods and spools. The original speed at Portland was 1.0/1.2s. Stroke times at Portland were increased by November 1991 in order to stop fatigue failure of components. The new stroke time at Portland was 1.6s (S.D. 0.3s) downstroke and 1.9s (S.D. 0.5s) upstroke for tests on 170 feeders. Both sets of stroke times were tested off- site. In August 1991, stroke times at T6 and Til were measured at 0.6/0.9s and 0.8/0.8s respectively. These stroke times were used for appropriate runs of T6 and Tl 1 feeder designs. Stroke times were set by flow control valves and a stop watch for off- site tests on the test rig.
2.3.4 Validity of Off-Site Testing.
It is far easier to do shot size testing off-site on cold pots than on hot pots where there are problems with magnetic effects, crane access, heat/dust/fumes and clashes with normal operations requirements. For these reasons, the majority of Portland tests were conducted on a test rig off-site. As outlined below, the validity of off-site testing was given careful consideration.
(a) Stroke Times Dwell time is the total of the following:
(i) wait time from DCV movement to when the stroke of the piston starts, (ii) stroke time, and (iii) waiting time when the piston is fully extended to when the DCV is reversed.
As discussed in Chapter 4, the wait times and the time to reach maximum pressure are much longer for on-site multi-feeder operation compared to one feeder operating with CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 70 its own DCV. Stroke movement time has the greatest effect on shot size accuracy with integrated feeders. For the tests off-site, stroke time was fixed to be similar to on-site operation and the dwell time was varied across a 0.5s to 5.0s range.
(b) Validity of Alumina Size Variations
It is possible that the repeated use of the same alumina for off-site testing could have given false results due to dust loss. To replace the alumina for each test would have required about 40t of alumina which was not acceptable. Alumina size did in fact increase as the runs progressed (Appendix 3).
To minimize the effect of dust loss, runs were conducted in blocks of about 4 runs. In addition, one feeder (Parker #5001 cylinder) using normal sized alumina and standard components was used repeatedly throughout the investigation to check if there was a significant change to the testing method. Results showed shot size changed when the alumina changed, but the feeder combination retained many of its shot size characteristics (Figure 2-4). Note that Runs 30 and 31 are not sequential due to a mix up in numbering.
For testing of the A3 feeder off-site, a worst case analysis was conducted. It was decided to thoroughly test the A3 feeder in the toughest manner possible in order to draw out any failures of what appeared to be the best dosing unit design. For this testing, alumina was replaced after each run. The alumina was taken from reservoirs in the superstructure which are usually the finest alumina used in pots. As will be seen, the A3 still proved superior in shot size accuracy.
(c) Temperature The temperature of alumina exiting the superstructure is about 60-120 C, as measured at Portland and T16. Arnold (44) tested alumina at ambient temperature and 100 C. He found that there was no discernible effect on parameters pertaining to the storage or flow characteristics of reacted alumina. Hence, tests at room temperature should exhibit similar characteristics to those expected at the elevated temperatures in a pot. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 71
FIGURE 2 - 4
PARKER STANDARD CHECKS #2 AEDD, NORMAL ALUMINA, NO INSERT )
SHOT SIZE (g) 2000
1800 y'y'" 91 ***"••**'.' .-•*' .-**•>.* — 1600 / /' / y 1400 / y' jr / / / m M 1200 /jf 1000 STROKE TIMES
7 RUN 22 1.0/1.1 800 RUN 31 1.2/1.7 RUN 30 1.7/1.9 600 —
— 400 RUN 22, JAN29 RUN31.FEB25 RUN 30, FE828
1 1 I L ... I 1— 200 _i 1 J L 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 DWELL TIME (s) STD DEV {%)
«. ""•nai»'*"*"*,*"»»». _L i RUNS 22,30,31 Note-. 1.7/1.9 refers to 1 Js downtroke and 1.9s upstroke time. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 72
This assumption proved valid in testing on cold pots and on operating pots (refer Sections 2.6.2 and 2.6.3).
(d) Times and Masses
Dwell time was controlled by an electronic timer. Stroke times were checked using a stop watch. A cross check was carried out with the stroke testing unit using a high resolution chart recorder. Times were accurate to +/-0.05s and were only quoted in 0. Is increments. Scales were checked daily with five test weights in case the instrument drifted. The masses were within +/-0.5g. Though shotsize were only quoted in lg increments.
(e) Continuity of Procedure
Standard procedures were carried out for all testing. Periodic checks revealed the procedures were followed consistently and on-site testing verified results. Only three people conducted shot size tests so there was minimal person-to-person variability. It was concluded that operating practices were well controlled, especially when one considers the boring and repetitive nature of the job.
(f) Unreacted versus Reacted Alumina The unreacted alumina behaved similarly to "normal" reacted alumina for all feeders (Figure 2-2). This was due to the fact that it was the same size as the normal alumina (Appendix 3). This agrees with the findings of Arnold (44) that they have similar bulk flow properties. To avoid any possible undetected effect on flow properties, reacted alumina was used for all but thefirst few runs.
Note that the above tests showed that feeding unreacted and reacted alumina gives similar mass delivery. However, this does not mean that the dissolution in the bath is necessarily similar. This is a possible shortcoming of bench scale testing of aluminas published to date (refer Section 1.3.13). CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 73
2.4 OFF-SITE TESTING OF SHOT SIZE
Data from all tests are provided in Appendix 2, and a summary of results is given in Table 2-1. Dwell time of 2.5s was used for comparison purposes. A more detailed analysis of this data is in Kissane (76). The key results were as follows:
TABLE 2 -1
SUMMARY RESULTS OF OFF-SITE SHOT SIZE TESTS
STROKE WITHOUT WITH FEEDER TIME INSERT INSERT (down/up) (s) (S.D.%) (S.D.%) PORTLAND AEDD 1.0/1.1 7-14 2 PORTLAND AEDD 1.6/1.9 2-10 1 T6 0.6/0.8 1-2 1 T11 0.8/0.8 1-3 <1 A2 1.6/1.9 <1 n/a A3 ALL <1 n/a
Notes: (i) Dwell time of 2.5s was used for comparison. (ii) "n/a" means not applicable
2.4.1 Portland AEDD
Consider Runs 1-27 and Figures 2-2 and 2-5. 1. The main influences on shot-to-shot mass changes were as follows: Different feeders 13% Alumina size 13% Type of cylinder 9% Dwell time 7% Stroke time 2% Overall effect is variation of up to 37% for any combination. Parker cylinders had the worst results. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE74
FIGURE 2 - 5
EFFECT OF CYLINDER DESIGN ON PORTLAND AEDD SHOT SIZE NORMAL ALUMINA, STANDARD SPOOL
SHOT SIZE (g) 2000
...... ».» »•» •»• ... • 1800 -MW. •*: Jf-> "*• — - -j-^rajpfc*!..*-
1600 '/ //
/•' // '•: ff 1400 Is " / 1200 * // '/ // f 1000 '' - x
800 "yh
600 - PARKER, 1.7/1.9 PARKER, 1.2/1.7 TERRY, 1.7/1.9 TERRY, 1.2/1.7 400
200 I 1 1 1 1 1 1 _| 1 1— 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 DWELL TIME (s) STD DEV (%) 6 5 4 3 ••>•*, 2
1 V _ .IL —.- —.-..—-»~?HI^'-«2!
0 RUNS 30,31,32,33 Note: 1.7/1.9 refers to 1.7s downstroke and 1,9s upstroke time. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 75 2. For fine alumina, the main factors were: Dwell time 13% Type of cylinder 7% Different feeders 2% Stroke time 2% 3. Slowing down the feeder to 1.6s downstroke had the following main effects: Cylinder type 7% Dwell time 7% for Parker <1% for Terry Alumina size 4% 4. Dwell times less than the plant stroke times caused smaller shot size and significant increase in variability compared to dwell times over the plant stroke times. 5. Six (6) feeders taken at random from off line pots were tested to see if the feeders being tested were representative of normal feeders (Figure 2-6). This showed similar results to the test feeders and confirmed that Parker cylinders were more affected by dwell time than Terry cylinders. 6. Shot size varied with several parameters, but was generally about l,600g at 2.0s dwell time for a standard AEDD feeder. 7. Initial tests of spool insert shape used cast iron inserts with and without enamel coating. Enamel coating had little effect on variability. As steel was cheaper to manufacture and cheaper to install, steel inserts were used for all future testing after Run 27 and later installed in the smelter in production designs. In the light of experience, there has been no deterioration of the surface of the inserts in production over a period of three years; if anything, steel insert surfaces became smoother over time. 8. For feeders with inserts, the stroke length had an effect on the magnitude of the shot size. There was drop of 17% in shot size for a 81mm stroke and 23% for a 53mm stroke compared to the same feeder without an insert (Figure 2-7). Standard deviations reduced from 7-14% to 2%. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 76
FIGURE 2 - 6
PORTLAND AEDD REPEATABILITY EX PLANT STANDARD AEDD, NORMAL ALUMINA, 1,6s /1.9s
SHOT SIZE (g) 2000
600 PARKER.1029 PARKER.1975 PARKER.1282 TERRY,580 TERRY.474 TERRY.650 a s & m a » »•» 400
200 U 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4,5 5.0 DWELL TIME (s) STD DEV (%) 6
RUNS 40,41,42,43,44,45 Note: 1.7/1.9 refers to 1.7s downtroke and 1.9s upstroke time. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 77
FIGURE 2 - 7
EFFECT OF INSERT STROKE ON PORTLAND AEDD SHOT SIZE PARKER CYLINDER, FINE ALUMINA
SHOT SIZE fg) 2000
1800
STANDARD 63mm STROKE 81mm STROKE «)•«•«* 400
200 -i 1 1 r- -i r- .5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 DWELL TIME (s) STD DEV (%) 6
RUNS 102,107,112 CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 78
9. For most runs, the 25 consecutive points were statistically "in control" if dwell time was greater than the stroke time as shown in Appendix 1 where the data falls between the UCL (upper control limit) and LCL (lower control limit) for both shotsize and range.
2.4.2 T6 Integrated Feeder
Consider Runs 1-23 and Figure 2-8.
1. Changes in dwell time and feeders had little effect on shot size i.e. about 1%. 2. Changing from normal to fine alumina dropped shot size by 6% in the first runs conducted, where there was a large difference in size between fine and normal alumina. Later runs showed little drop in shot size as the size difference reduced (Appendix 2, Runs 50-53 and 72-74). 3. Fine and normal alumina developed similarly sloped curves. 4. The S.D. was less than 2% of shot mass for a standard feeder and 1% if the
feeder had an insert. 5. Shot size was about l,700g at 0.6s stroke time for the standard feeder and
l,300g if it had an insert; a drop of 24%. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 79
FIGURE 2 - 8
T6 FEEDER SHOT SIZE 53mm SPOOL STROKE
SHOT SIZE (g) 2000
i,t , 1800 »:*W».3»8*»* »*'» *' .».w».[[| »»-—•»»„ 1800
1400
.tf •*• «*•*-"•*• «»•*•.«•• -m. "I..™ Jd *J! 1200
1000
800
NORMAL.WITHOUT.0.6/0.9 RNE,WJTHOUT.0.6/0.9 NORMAL,WiTH50.6/0.9 600 h
FINE,WITH,0.6/0.9 NORMAL,W!THOUT,1.1/1.3 FINE,WITHOUTS1.1/1.3 400
200 l-i 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 DWELL TIME (s) STD DEV (%) 6 V '; 1 " \ . 1 5 — i, • i ~ * • 1 4 - \. I V 1 ? 3 - a 1 S VI " ," * 2 —N. "• / 1 *• * * * - *" " >--a-*-^S- - 0 _l 1- !• 1 1 1 1 1 1 '—' RUNS 50,51,52,53,54,55 Note: 1.7/1.9 refers to 1.7s downtroke and 1.9s upstroke time. CHAPTER 2 SHOT SIZE ANALYSIS PAGE 80 2.4.3 Til Integrated Feeder
Consider runs 56 to 61 and Figure 2-9 and 2-10.
1. The two feeders tested operated quite differently. One operated like the Portland AEDD with a Parker cylinder and the other like the Portland AEDD with a Terry cylinder. Dwell time had little effect, but one feeder (#2) was badly affected by dwell time. Most testing was done on #2 to simulate the worst case feeder.
2. Changing from coarse to fine alumina dropped shot size by 4%. This was reduced to 1% with the insert. 3. Shot size was similar for both feeders for fine alumina, but not for normal alumina over a dwell time of 2s.
4. The insert installed in #2 feeder brought all shots to a S.D. within 1% of shot mass for dwell and alumina size changes versus 2% without the insert. 5. Shot size was identical for normal andfine alumin a with the insert installed, but the size dropped from about l,600g to l,300g viz. 25%.
2.4.4 Portland A2 Feeder
Consider runs 4 to 64 and Figure 2-11.
1. Shot size changed by less than 1% for changes in dwell time (at fast and slow speeds) and alumina size. Hence, the feeder was not affected by the alumina (presumably due to the 60 angles at the top and bottom of the dosing unit which assist mass flow). 2. Stroke time affected shot size by over 4% for a 0.5s stroke time change. 3. Cylinder type changed shot size by 2%. 4. Shot size was 900 gram (virtually independent of dwell time). 5. S.D. was similar to the AEDD with insert viz. 2%. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 81
FIGURE 2 - 9
Til FEEDER SHOT SIZE
SHOT SIZE (g) 2000 - 1800 - 1600 ^ *t»v**»"** ^ •"* r , » f > a» •*" 1400 A* ' "'
1200 - /' ' ' - .- -ft/ / 1000 ~ 1 _ / 800 _ ? f - f 600 #1, NORMAL, 0.8/0.8 #1, FINE. 0.8/0.8 #2, NORMAL, 0.8/0.8 400 #2, FINE, 0.8/0.8 #2, NORMAL 1.3/1.3 #2, FINE, 1.3/1.3 200 I I I U I I I . I I I 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5. DWELL TIME (s) STD DEV (%) 6 It 5 4 3 - *Vl
2 \» »•• "" Ha
1 \ s. _ _ _ _ _ „, ** »•• .*
0 I • 1 1 1 1 1 1 a 1 1 1 RUNS 56,57,58,59,60,61 Note: 1.7/1.9 refers to 1.7s downtroke and 1.9s upstroke time. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 82
FIGURE 2 -10
Til #2 FEEDER WITH AND WITHOUT INSERT
SHOT SIZE (g) 2000 1800 r
NORMAL FINE NORMAL WITH INSERT FINE WITH INSERT
• »« •••
200 L-1 • •- H H H 1 1- 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.I DWELL TIME (s) STD DEV (%) 6 5 4 3 2 1 0 RUNS 58,59,62B,62A CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 83
FIGURE 2-11
A2 FEEDER SHOT SIZE
SHOT SIZE (g) 2000
1800-
1600
1400
1200
1000 ...... ^• 600 ft PARKER.NORMAU1.0/1.1 PARKER,F1N£,1.Q/1.1 TERRY,NORM AL.1,7/1.9 ••HI^aB RSR»R«XV »•• mm* ««« > 400 TERRY.FINE.1.7/1.9 PARKER!NORWIAL:1.2/1.* ARKER.NORWIAL,1.7/1.5 200 "-1 u x 4- 4- 4- 4- 4- 4- 4- 0.5 1.0 1.5 2,0 2.5 3,0 3.5 4.0 4.5 5.0 DWELL TIME (s) STD DEV (%) 6 af I4** $••?,$ %7-.m- ««-.* RUNS 10,4,65,67,63,64 Note: 1.7/1.9 refers to 1.7s downtroke and 1.9s upstroke time. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 84 2.4.5 A3 Feeder and Sequential Feed As the A3 and sequential feed designs were not affected by dwell time, graphs as shown in Figures 2-4 to 2-11 are not relevant. Table 2-H shows a comparison of the AEDD , A3 and integrated feeders. Shot size testing has been carried out on nine A3 feeders. To test sequential feed operation, a Portland AEDD (with insert) was tested with a DCV mounted on the top flange. (Previous off-site testing had one feeder per DCV but the valve was mounted about 2 metres away from the feeder.) TABLE 2 - II A3/AEDD SHOT SIZE COMPARISON (ROOM TEMPERATURE) Feeder Mean(g) Feeder-to-Feeder Shot-to-Shot Number Location S.D.(g) S.D.(%) S.D.(g) S.D.(%) of tests AEDD 1223 69 5.7 30 2.4 145 on-site A3 1276 17 1.3 6 0.6 9 off-site SEQUENTIAL 1270 7 0.5 1 off-site A3/AEDD 25% 23% 20% 25% Notes: 0) AEDD data from new pots testing (Table 2-IV) of AEDD feeders with inserts. (ii) A3 from slowest dosing cylinder stroke speed off-site (10 shots). (iii) Sequential from 2s dwell time off-site (10 shots). Observations from the tests were as follows: 1. Standard deviation for the A3 feeders did not change over the complete range of speeds possible from the DCV. S.D. was 0.6% of shot mass. 2. Shot size for the A3 feeders was about l,280g, but shot size reduced by 40g at maximum speed compared to slowest speed. As fatigue and wear of parts is minimized at the slowest speed, this setting was selected to provide maximum feeder life for plant operation. 3. Shot size and S.D. for the sequential feeder were similar to the A3. 4. The type of DCV (Mac or Atlas Copco) had no effect, nor did the type of feed cylinder (Ortman, Parker, Kempe, Terry). This suggests that any of the hardware on trial can be interchanged and not affect the feeder performance. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 85 Clearly, the A3 is superior to the AEDD with respect to feeder-to-feeder and shot-to-shot shot size accuracy viz. four times better. These results may be typical of other independent feeders as the DCVs and relative feed cylinder area of the A3 are similar to that of Pechiney and D3 pots. 5. Note that the dosing assembly for the sequential feeder was the same as that used on all the AEDD tests (#2). The air feed control system was different but the dosing unit shape was identical. It was concluded that the main reason for poor AEDD feeder accuracy compared to sequential or independent feed is the pneumatics of the pot rather than the feeder dosing shape. 2.4.6 Decision on Use of Spool Inserts The main attraction of the insert is to minimize the effect of alumina particle size. If segregation is a problem in a plant, then the plant should consider the benefits of redesigning the dosing unit shape. For T6 and Til feeders, shot size variability in almost all runs improved with an insert (generally to half the standard deviation), however, shot size reduced by 15-20%. This requires more air usage, more cycles of the feeder and more time that the plunger is wet. As shotsize accuracy is better than 2-3%, it is debatable if the reduction to under 1% is worth the high cost of installation and maintenance due to increased cycles. Note that the two Til feeders tested at Portland reacted quite differently. Maybe the same could be said if other T6 and Til feeders were tested as well. Other factors such as cylinder, spool, and spring quality can affect shot size; the insert minimizes the effects of these. Due to the relatively good shot size accuracy of the T6 feeder and the difficulties of shot size reduction for a plant which has a history of difficulties with lack of air, T6 decided not to install inserts. For Portland, the benefits of inserts are quite clear and justifiable from off-site testing results. The perceived benefits to pot operation were considered to be worthwhile for the increased cycles required for a 25% drop in shotsize. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 86 For independent feeders, the increase in air usage would have little effect on plant usage as the feed cylinder is small. Angles of 60° at top or bottom of independent feeder dosing units is recommended rather than the 45°commonly used. As with T6 and Til, the improvement is probably not economic for a retrofit, but probably worthwhile on new feeders in order to reduce variability for virtually no difference in cost for a new part. From May 1992, inserts were installed in Portland feeders with 81mm stroke. However, it was found that in almost all cases a new spool and spring was required. This took a $40 change-over to $240 which was too expensive. Further testing was carried out on stroke length (refer to Section 2.5.7) and the stoke was changed back to the original 53mm stroke as of September 1992. By June 1993, plant data confirmed the benefits of the insert shape on pot performance. Table 2-JJI shows that feeders with inserts had reduced anode effects by up to 17%. Multiple anode effects per day (AEPD) by up to 31% and S.D. within a pot reduced by up to 15%. Also, it appeared that the longer the insert was installed, the better the results became as feeders installed by over 60 days gave better results than pots over 30 days old. This was due to better tuning of the computer interval between shots. Hence, from July 1993 a program of rapid feeder changeout was employed to implement this change as soon as possible. TABLE 2 - ffl ANODE EFFECTS FOR POTS WITH/WITHOUT SPOOL INSERTS MONTH IN WHICH DAYS NUMBER OF UNSCEDULED AEPD MULTIPLE AEPD STD DEV WITHIN A POT POT WAS STARTED IN POT WITHOUT WITH CHANGE CHANGE (%) WITHOUT WITH CHANGE CHANGE (%) WITHOUT WITH CHANGE CHANGE (%) JANUARY >30 0.63 0.62 -0.01 0.13 0.12 -0.01 0.82 0.80 -0.02 I >60 0.63 0 53 -0 10 0.13 0.11 -0.02 0.81 0.75 -0.06 FEBRUARY >30 062 0.61 -0.01 0.12 0.12 0.00 0% 0.78 0.74 -0.04 5% >60 0.62 0.53 -0.09 0.13 0.09 -0.04 31% 0.77 0.67 -0 10 13% MARCH >30 0.64 0.58 -0.06 I 0.13 0.11 -0.02 I 0.78 0.70 -0.08 10% >60 0.63 0.52 -0.11 0.13 0.09 -0.04 0.78 0.66 -0.12 15% Notes: (a) pots started after July 1992 had Inserts ("WITH") and pots with 8kg shot size did not ("WITHOUT"). (b) values listed in 'WITH" and "WITHOUT" refers to average variability within a pot over the period selected. (c) "CHANGE" refersto "WITHOUT" minus "WITH" values for >60 days worth of data. (d)this analysis was calculated on 27th April 1993. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 87 By the end of 1993, virtually all Portland feeders had spool inserts and anode effects were at record low levels. Since then, anode effects have fallen still further. The inserts were an enabler to anode effect reduction but were not the sole reason for the improvement due to considerable attention to better computer control and greater awareness of the importance of reduced anode effect on pot performance. Computer control could be fine tuned, knowing that the feeder shot size was now reliable. 2.5 SENSITIVITY ANALYSIS ON SHOT SIZE As noted in Section 2.4, the AEDD feeder design had quite varying results from feeder- to-feeder and plant-to-plant. A sensitivity analysis was conducted to identify which parameters affected feeder accuracy the most, so that design changes could be made to minimize the variability. 2.5.1 Bulk Density of Alumina Analyses of the bulk density of alumina (loose and packed) for different sizes showed little significant change (Appendix 4). The average loose bulk density was 1,028 kg/m3 (S.D. 5 kg/m3) for normal reacted alumina and 1,044 kg/m3 (S.D. 13 kg/m3) for fine reacted alumina. At a 90% confidence level, there is no difference between standard deviations of fines versus normal alumina for either loose or packed bulk densities. However, the means are statistically different for each type of density. As there was no vibration or any other disturbance to the hopper of the off-site test rig during the tests (and little in an operating feeder) there should be no change to the extent of compaction from one run to another. Hence, the density should be consistent for all runs. It seems reasonable that loose bulk density is the more suitable estimate of the density in an operating feeder. If so, one can see that the difference between fine and normal reacted alumina is small (1.5%). CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 88 Hence, it was concluded that any change in shot size greater than 1.5% for different feeder designs in the tests is a reflection of dynamic factors such as: (i) the angle of flow into and out of the dosing unit; (ii) the angle of repose in the dosing unit; and (iii) the amount of material left in the dosing unit at the expiration of the dwell time. For almost all variables tested, the differences seen were greater than 1.5% so they were more important to shot size accuracy than bulk density. It is also postulated that a volumetric feeder is an adequate meter for feeding fixed masses of alumina to a pot. There appears to be little requirement to explore gravimetric feeders. Shapiro (71) investigated such a device in 1981-83. This device achieved standard deviations of 1.3% - 3.6%. Testing in Portland has identified standard deviations of under 1.0% are achievable for AEDD feeders with inserts (Table 2-1) and standard AEDD feeders at T6 and Til achieved this without inserts. Hence, the significant extra cost of a gravimetric feeder is not justified. 2.5.2 Flow Through Inaccuracies in Dosing Systems The spring in a Alcoa feeder pushes the spool downwards when the cylinder extends (Figure 1-9). It must do this as fast as possible so as to minimize unmetered flow of alumina through the spool. Once the cylinder starts to move down, immediately the alumina starts by-passing through the spool until the spool stops. If flow through occurs, this unmetered alumina adds to the variability of the shot size. As flow is affected by alumina size and degree of aeration, spool movement in an AEDD design is affected by the quality of cylinder and spring (which always touches the spool). For independent feeders, flow through also occurs, but the spool is controlled by the small pneumatic feed cylinder, cylinder design and air supply. As the DCV is mounted close to the cylinder, and there is only one DCV per feeder, there is CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 89 more repeatable control of piston movement. In both types of feeder, the speed of the spool affects shot size accuracy by varying the amount of flow through. 2.5.3 Effect of Springs Shot size testing was conducted with various spring configurations (Runs 113-129). The observations from this testing included: (i) the spring condition affects the standard AEDD shot size. (ii) high pressure (Kempe) springs make the feeder less sensitive to cushion quality. (iii)spools with inserts are less sensitive to spring condition. All Alcoa, Alcan and Alumax integrated feeders until 1993 used 302SS or 304SS springs of identical shape (until this study recommended to a change to high pressure steel springs). The normal free standing height of the standard AEDD stainless spring reduces due to creep from exposure to heat (Figure 2-12). As the spring reduces in length, it reduces its spring pressure which in turn changes the characteristics of the spool movement. A soft spring will not act as well as a brake for the spool on the upstroke, and will not force the spool down as hard on the downstroke. Less braking on the upstroke gives less bypass and better shotsize variability. The net effect of the slower downstroke speed (and deceleration) is a tendency to retain alumina. This causes a slow dribble of alumina to flow out, so the shot size is affected by dwell time and the variability increases. A sharp downwards action tends to throw the shot out of the feeder giving less variability. The downstroke effect is considered to be more important than the upstroke as witnessed by the effect of varying dwell time on shotsize. The Kempe spring has been designed to cope with higher temperatures and to exert high forces at the extended and compressed lengths it is exposed to in the feeder. Its force does not drop to the original value of a new stainless spring until it reaches over CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 90 500 C. The stainless springs start to lose strength at 250°C and to shrink in length at 300 C. Spring temperature reaches well over 300°C on some feeders (Section 6.5.2 and 8.4.2:2). FIGURE 2 -12 EFFECT OF TEMPERATURE ON SPRINGS 300 r 600 p - KEMPE HEIGHT i - STAINLESS HEIGHT — KEMPE FORCE — STAINLESS FORCE 25 200 250 300 350 400 450 TEMPERATURE (°C) Note: Springs were subjected to 18hrs for each 50°C temperature increments Force tests on old and new springs are illustrated in Table 2-iy From this table, the following observations can be made: (i) When the spool is fully down (cylinder extended), stainless springs drop in force by 60% after they have shrunk by 40mm. A reduction of 40mm is common in many plants. (ii) The closing force of stainless springs, even when new, is very small viz. 40- 120N. It does not take much restriction by foreign objects or alumina to prevent the spool closing off flow from the superstructure. This can lead to by-pass of alumina into the pot on a continuous basis during the wet time. This alumina is not measured and affects shot-to-shot variability. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 91 (iii) Kempe springs exert a 150% higher force when compressed than a new stainless spring. However, the force is much lower for both springs than the force exerted by the cylinder, so this does not affect the speed significantly. Hence, on the upstroke, the spring does not affect alumina bypass significantly. (iv) The closing force of the Kempe spring when the cylinder is extended is about double that of the stainless spring. (v) Use of length as a criterion to reject springs (which is often done in repair shops) is unwise. One must use force, as the force can drop up to 20% before the length reduces. Portland started installing Kempe springs in July 1991 and has installed about 1,300 to July 1994. (Refer Section 8.8.3 for discussion on spring failures) 2.5.4 Effects of Cylinder Design on Shotsize 2.5.4:1 Description of Cylinder Operation From Section 2.4, one can see different cylinders (Parker, Terry, Lindberg) had different effects on shot size of integrated feeders. This was due to cushioning, seal type and seal condition. These factors are discussed below, but let us first consider cylinder cushioning. Figure 1-20 shows the critical parts of a typical pneumatic y cylinder and Figure 2-13 shows how cushioning occurs. A pneumatic cylinder operates by application of air to alternating sides of a piston in a barrel. Cylinder manufacturers design cushioning based on the calculated kinetic energy of the mass being driven and the anticipated compression of the air to stop the piston in a controlled manner. The clearances between the piston rod cushion bushes and the fixed orifices of the cylinder mounting plates dictates the cushioning ability of a given cylinder design; the closer the clearances, the better the cushioning. For many cylinder designs, there are cushion adjustment screws to throttle the rate that air is exhausted from the cylinder. One can increase cushioning by restricting exhaust air flow. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 92 FIGURE 2 -13 CYLINDER CUSHIONING OPERATION CUSHION SPEAR CUSHION BUSH CUSHION ADJUSTMENT SCREW The main* air supply to the rur and of the cylinder is rapidly distributed over the'entire piston face with the aid of a non-return valve, built into the cylinder end cover. This drive* the piston on the out-stroke, and air forward of the piston is exhausted through the front port until -die seal of the cushion boa* enters the forward recast The remaining trapped air then escapes slowly through ai adjustable needle valve, thus providing a cushion affect ai the and of the piston stroke. On the reverse operation, (in-stroke), the cushion effect la achieved in the same manner. Adapted from Martonaire CIO 2.2.69-17 Brochure, Pg 17,1.4.77 If cushioning is poor, piston rods or pistons can break due to excessive deceleration rates. This has occurred at Portland, T6 and T17 (Section 8.6 and 8.8). The presence of flow control valves at most plants helps correct for poor cushioning settings during maintenance. These allow regulation of back pressure air to the CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 93 cylinders to control the speed. One will find that most operators or tradespersons who look after feeders are aware that changing the settings on these flow control valves stop the feeders from "banging". This is due to inefficient cushioning. At the above mentioned plants which have piston rod failures, there are no flow control valves so the cushioning may be ineffective. Clearly, there is a link between cushioning and fatigue. In most situations, cushioning efficiency has not been perceived to be a problem for maintenance, probably because cushioning is difficult to test in a workshop. Hence, cushioning has not been examined in any detail. However, feeder cylinders have a significant effect on shot size which few people (if any) have examined to date. This is more important than maintenance, as this affects pot performance. In mid 1991 a cushion test rig was installed at the Portland Kempe workshop to test all feeders after they were overhauled to ensure cushioning was adequate. A similar device (though not as good a design) was installed at Boyne in 1993. Ensure that any cushion test equipment is extremely anchored and that there is zero movement in the frame when a feeder is tested. If the frame has any flexibility, it will affect the accuracy of the measurement. This was discovered on the Parker production test rigs where the deceleration for cylinders increased by a factor of 10 when cushion feet were removed. This can be assured by installing (for a once off test) an accelerometer on the frame. 2.5.4:2 Effect of Cushioning on Shot Size Independent feeders have the spool driven by only a very small cylinder (40-50mm O.D. x 40-50mm stroke) that solely operates alumina dosing. Integrated feeders are driven by a much larger cylinder (100-150mm O.D. x 300-550mm stroke) but the piston rod has to drive not only the dosing unit but also the plunger shaft and plunger. The driven masses are about 10kg for independent feeders and about 30-50kg for integrated feeders. In order to cushion the larger driven mass, integrated feeder cylinders need much more extensive cushioning otherwise the piston will impact on the stationary metal stops and result in fatigue of piston and piston rod. Unfortunately, this larger cushion has a deleterious effect on shot size accuracy. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 94 For integrated feeders, as the cushioning is engaged, the speed of the driven mass is affected. Any change in speed in the cushioning area of the cylinder is directly related to the spool speed as the piston rod is directly coupled to the spool of the dosing unit. Hence, the speed of the dosing spool is affected, especially as the spool starts to move and when it slows to a stop. For the AEDD integrated feeder design, the piston is attached (by the piston rod) to the plunger shaft which has a collar just above the plunger (Figure 1-11). When the piston is about 50mm from the end of travel on the upstroke, the collar strikes the spool at the bottom surface (the cup). The piston spear hits the cushion sleeve about 25mm from the end of travel for integrated feeders (Figure 1-20 and 2-13). Then the piston slows down progressively to a stop just before it hits the end of stroke. Hence, the cup is struck before the cushion is engaged at a high speed...about 0.4-1.4 m/s (Table 3-1). Independent feeders have much smaller cushion lengths (10-20mm) and the velocity is much smaller (due to the short stroke length), so the cushioning effect on speeds is less. In both independent and integrated feeder dosing units, alumina can flow unmetered through the dosing spool as the cushions are engaged. The longer the cushioning, the potential for by-pass is worse. Thus, the longer the cushioning length or the tighter the air clearances between moving and stationary cylinder parts, the worse is the shot size variability. Testing of the two Til feeders showed that the one with the tighter cushioning (#2) had worse shot size variability than the one with sloppier cushioning (#1). Refer to Runs 56-61 and Section 2.9. The "best" cylinder design (i.e. the one which is least likely to fatigue) is one with good cushioning. However, this is expected to give the worst shot size variability. Shot size tests for Portland AEDD (Figure 2-14) verified that having no cushioning at all on a cylinder gave more consistent shot size and had less variability than the installation of an insert. Inspection of cushioning by the author at 13 plants during 1991-93 showed a variety of clearances. In the case of cylinders at Portland (Terry cylinders), T6 and T15, there was no cushion adjustment screw, so speed could be not CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 95 FIGURE 2 -14 EFFECT OF CUSHIONING ON PORTLAND AEDD SHOT SIZE PARKER, 1.6s/1.9s SHOT SIZE (g) 2000 »:CS*IWK!!.*l usssa. asyrts^aitftvcuf* STANDARD S3 INSERT STANDARD, NO CUSHION S3 INSERT, NO CUSHION RiPVntf if SV • *• •• 0.5 1.0 1.5 2.0 2,5 3.0 3.5 4.0 4.5 5.0 DWELL TIME (s) STD DEV (%) 6 RUNS 129,127,123A,124 Note: 1.7/1.9 refers to 1 Js downstroke and 1.9s upstroke time. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 96 be adjusted. The cushioning was poor. Lindberg cylinders at Tl 1 showed uneven wear on the cushion spear of the two examined at Portland. The one with the most wear (#2) was the one with the worst shot size results. To minimize the by-pass effect in dosing systems of integrated feeders, the spool insert shape was designed to incorporate a shoulder at the base so that the dosing unit hardly opens when the cylinder is in the cushions (Figures 1-11). Also, the steeper angles (60° versus 45 ) allow all the alumina to exit quickly from the dosing spool so there is little cushioning effect at top or bottom of the spool movement. This has proved to be successful as evidenced by discussion throughout this Chapter. 2.5.4:3 Effect of Seal Material/Quality Six feeders (three Terry and three Parker) from off-line pots showed a spread of shot size as illustrated in Figure 2-6. This illustrates the variation from feeder-to-feeder due to wear differences. A more detailed experiment was conducted on one feeder with old and new seals with and without inserts (Runs 130,131,132,134). This showed up to 10% greater shot size variation for old seals, but this reduces to under 2% if inserts are installed (Figure 2- 15). Thus, once again, inserts minimized variability. The type of piston and rod seals affect spool speed. These seals may be Teflon (as used in most Alcoa/Alcan/Alumax integrated feeder plants) or Viton (as used Atlas Copco and CPOAC cylinders at several independent feeder plants (Table 1-H), and, until recently, on Scheffer, Hanna, Terry and Parker cylinders at Portland, T8, T9, TIO, T12). Teflon is self lubricating and hence allows faster movement than Viton which is a rubber compound. Portland changed to Teflon rod seals and piston seals in 1992 for all Terry cylinders overhauled. Trials are underway on Teflon piston seals on Parker cylinders at Portland. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 97 FIGURE 2 -15 EFFECT OF SPOOL DESIGN & SEAL AGE NORMAL ALUMINA, PARKER #1248 EX PLANT, 1.6/1.9 ] SHOT SIZE (g) 2000 - 1800- ,/\. —-*^T— "" "• '" ^^^ —* 1600 / ^ — f »** 1400 *r f v********9***^ ' J / 1 ¥•"' 1200 ' ''/'' t y/ H'/J 1000— ¥tf/ w ^ - ,'/ f ft J 800 1 / '/ 600 _/ OLD SEALS,STANDARD SPOOL OLD SEALS WITH S3mm INSERT 400 NEW SEALS, STANDARD SPOOLNEW SEALS WITH 53mm INSERT — ••» • •• "^ *• ^ 200 lit 1 J 1 1 1 1 J- 0.5 1,0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 DWELL TIME |s) STD DEV {%) RUNS 130,131,132,134 Note: 1.7/1.9 refers to 1.7s downtroke and 1.9s upstroke time. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 98 The flexibility of the material can have a significant effect. For example, the Viton used for Parker cylinders is a hard material, but that used by Atlas Copco is very soft so it can take a higher temperature excursion before becoming brittle. The quality of the seals can also affect speed. If there is air by-passing the rod seals, the effective air pressure on the bottom face of the piston is less, resulting in reduced piston speed if the manifold air is limited. Note that rod seal failure is one of the most frequent failures of feeders surveyed (as evidenced by ore leaks in Table 1-V). Piston seals are one of the most common causes of failure of crustbreaker cylinders on independent feeders and are probably just as bad for Alcoa and Alcan pots (but they have not been monitored in most plants). Reports by tradespersons at all eight AEDD plants visited by the author in 1992 suggest a very high piston seal failure frequency. A cylinder that has recently been overhauled will be slower than an old one (due to tighter clearances on seals and bushes) and this consequently affects alumina bypass through the integrated feeder spool. Shotsize testing of new and used feeders on start up pots at Portland (Table 2-TV) has shown that old feeders have a slightly more accurate shotsize compared to new feeders viz. 50g S.D. versus 69g S.D. with inserts (or 69g S.D. versus 147g S.D. for no inserts). This is expected to be due to wearing of seals and increasing the velocity of the dosing unit. TABLE 2 - IV SPRING FORCE COMPARISON RUN SPRING NOMINAL ACTUAL NEW FORCE (N) TYPE FREE FREE OR SPOOL STROKE FROM TOP (mm) HEIGHT (mm HEIGHT (mm USED CLOSE 10 20 30 40 OPEN A 304SS 260 250 NEW 206 191 171 152 137 118 B 304SS 260 250 NEW 211 191 176 157 137 118 C 304SS 260 210 USED 132 113 93 69 54 39 D 304SS 260 215 USED 144 127 108 88 69 54 E KEMPE 190 180 NEW 700 600 500 420 330 260 F KEMPE 190 177 USED 554 470 392 314 240 171 G KEMPE 210 210 NEW 583 500 426 348 279 196 Notes: (i) "Closed" height is equivalent to when plunger is fully retracted (130mm spring height). (ii) "Open" height is equivalent to when the plunger is fully extended (180mm spring height). (iii) "Spool stroke fromtop" i s distance in mm that the spring is extended from a closed height. (iv) Forces for Kempe runs E&F were using 30mm spacer with the spring. CHAPTER 2 SHOT SIZE ANALYSIS PAGE 99 2.5.5 Effect of Dwell Time Section 2.3.1 showed that increasing dwell time increases the shot size on most Portland feeders. One would expect a change in shot frequency in the plant if the dwell time has changed on all pots. At Portland there are equal numbers of Terry and Parker cylinders. The shot size tests of six (6) random feeders (Figure 2-6 and runs 40-45), showed that for a change of dwell time from 2s to 4s the shot size increased from l,710g to l,795g; arise o f 5%. For a fixed current supply to a pot, the pot requires a fixed amount of alumina. If the shot size increases by 5%, then the shot frequency (which is automatically controlled by the potline computer) will increase by 5%. In 1992, the dwell time on all pots was increased from 2.5s to 4s as a result of slowing feeder speed to improve fatigue of piston rods and spools. It was found that the pots required about 3% more frequent feeds to maintain stable alumina concentration in the pots. Although the plant change was less than that expected, it was of similar magnitude. The difference between the 5% predicted and 3% actual drop was probably due to the different pneumatic controls for off-site and on-site feeders. 2.5.6 Effect of Height of Alumina in the Hopper Level of alumina in the hopper did not affect shot size from 1.4m to 0.6m above the dose unit (Appendix 1). No testing was done for alumina height beyond these heights because they relate to a Portland pot superstructure maximum and minimum alumina inventory respectively. 2.5.7 Effect of Insert Stroke Length The original spool stroke length is 53mm for most AEDD plants. The original spool insert design increased this to 81mm as the insert had a 30mm shoulder at the bottom end that may restrict the discharge (Figure 1-11 and 2-3). Inserts were installed using the 81mm stroke from May 1992. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 100 As noted in Section 2.4.6, 81mm insert stroke was too expensive and action was taken to test if 53 mm stroke was accurate, and what was the effect on shot size. Comparison of the two strokes in Figure 2-7 indicated that the 53mm shot size was (i) 23% smaller than the standard AEDD (versus 17% for 81mm stroke). (ii) more repeatable. (iii) less sensitive to alumina type . (iv) less sensitive to cylinder type. Hence, it was decided to change to a 53mm stroke, effective from September 2nd, 1992. By December 1993, inserts were installed on over 98% of Portland feeders. The 25% drop in shotsize predicted by the off-site testing was evidenced by a similar increase in feed frequency in the potlines. 2.5.8 Effect of Spool Manufacture Quality Investigation of the quality of the spool dosing unit at Portland in 1991 showed that the spools were not concentric. Similar results were observed in 1992 at T9 and Til. Figure 2-16 illustrates the Portland data on 60 random spools showing that 90% were not concentric. Non-concentric operation can result in intermittent or complete jamming of the spool in the assembly due to the tight clearances. The spools should be spun in a lathe to ensure they are built to original drawings if a new spool is installed in a AEDD feeder. At Portland, about 1,000 spools were made of larger diameter pipe that fouled in the assembly and had to be replaced. Analysis of the dynamics of operation of the spool also detected that in all AEDD feeders a partial vacuum forms in the spring chamber (Figure 1-11). When the spool moves downwards, a partial vacuum is caused due to the increase in volume in the spring chamber. The only way air can generally flow into the chamber is from the dosing unit itself. Alumina will also attempt to flow into the chamber through the tight clearances between spool and feeder assembly which can lead to jamming of the spool in the assembly causing: CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 101 (i) ore leaks, (ii) not feeding, or (iii) variable feed rates. This can result in variability in shot size and intermittent alumina dosing. This phenomenon will be generally undetected by operating personnel, as the feeder will appear to be working correctly (unless the spool is totally jammed). FIGURE 2 -16 SPOOL WEAR POINTS DISTRIBUTION OF WEAR POINT AT THE TOP OF THE SPOOL PIPE 3% RUN OUT POINT AVERAGE = 0.7 mm STD. DEV. = 0.4 mm no wear at all = 10% Note: (i) sample size: 60 spools taken at random ex plant (ii) measurements taken when spool in lathe (iii) tests carried out in January 1991 at Portland. (iv) "Distribution" refers to the percent of spoors tested that had wear at the points indicated at the top of the spool. About half the feeders removed at Portland for "not feeding" up to early 1994 were probably due to this cause as there was no mechanical fault with the feeder when the feeder was overhauled after service. From early 1994, these factors were addressed at Portland by drilling holes in the assembly to reduce the partial vacuum by 75% (Figure 1-11). Immediately the frequency of jammed objects dropped (Section 8.8). CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 102 2.6 ON-SITE TESTING OF SHOT SIZE 2.6.1 Til-ColdPots Plant Til carried out shot size testing of feeders with and without inserts in cold pots to verify the benefits (if any) of the insert design and to gain a better understanding of the accuracy of AEDD feeders (76). These tests on regular feeders confirmed the tests done at Portland on two feeders from Til. The shotsize for different feeders varied between l,640g and l,770g - an 8% change. Shot size was larger than Portland tests of l,500-l,690g. Comparison of results of in-plant testing at Til and Portland shows that Til had a better S.D. than standard Portland feeders viz. 1% versus 3% (76). This is probably related to the fact that Til has only 3 feeders/DCV (100 x 356mm cylinders) versus 5 feeders/DCV (125 x 508mm cylinders) at Portland (refer Section 4.6). Both plants have similarly sized DCVs so the Portland feeders are relatively starved of air. It is easier to start the Til feeder spools, so the spool is quicker to open and close at the start of the upstroke and the end of the downstroke. Benefits of the insert were not evident at Til probably as the original design had a good variability anyway due to less feeders/DCV. It is likely that other plants with 1-3 feeders/DCV would find the insert may not give a significant improvement to shot size accuracy. Thus, the effect on pot efficiency may be insignificant. 2.6.2 Portland - Cold Pots From August 1992, each feeder in new pots has had its shotsize tested prior to start-up. A sample of this data is given in Table 2-V. Using data from 229 feeders, the benefits of inserts are clear...61% improvement in standard deviation within a pot, and 37% improvement feeder-to-feeder for new feeders. Reused feeders showed a 38% improvement in standard deviation within a pot and a 20% effect feeder-to-feeder. The reason for the S.D. for all feeders having a larger value than that within a pot is due to the variability of pot-to-pot changes because of different DCV and air supply characteristics. This verified the off-site testing. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 103 TABLE 2-V SHOT SIZE TESTING ON NEW POTS NEW FEEDERS INSERTS NO INSERTS MASS RANGE CALCD STD DEV S.D. AS % MASS RANGE CALCD STD DEV S.D. AS % WITHIN POT OF MASS WITHIN POT OF MASS AVERAGE PER 1223 53 23 1.9 1661 190 82 4.9 POT STD DEV FOR ALL 69 5.6 147 8.8 FEEDERS NUMBER TESTED 145 19 IMPROVEMENT FOR FEEDERS WITH INSERTS: STD DEV WITHIN POT DROP IN SHOT SEE WITH INSERTS: I STD DEV FOR ALL FEEDERS REUSED FEEDERS MASS RANGE CALCD STD DEV S.D. AS % MASS RANGE CALCD STD DEV S.D. AS % WITHIN POT OF MASS WITHIN POT OF MASS AVERAGE PER 1239 46 20 1.6 1670 102 44 2.6 POT STD DEV FOR ALL 50 4.0 69 4.1 FEEDERS NUMBER TESTED 43 22 IMPROVEMENT FOR FEEDERS WITH INSERTS: STD DEV WITHIN POT DROP IN SHOT SIZE WITH INSERTS: ISTD DEV FOR ALL FEEDERS Notes: (i) All mass in g. (ii) "Range is the maximum minus minimum of individual shot sizesfor 5 shots from 5 feeders. (iii) Calculated standard deviation within a pot used conversion data from Janus (96) to convert a range to a S.D. (iv) S.D.for allfeeders is calculated as the pooled S.D. of all thefeeders in the sample size e.g.for newfeeders wit h inserts, the S.D. of each of 145feeders was squared and the square root of the average was 69g. This represents 5.7% of 1223g, which is the average mass taken from the 145 feeders. 2.6.3 Portland - Operating Pots There is always an element of doubt over room temperature tests as they may not be applicable at pot operating temperatures. To verify the effect of the inserts and to get a better feel of other factors affecting shot size at operating temperature, tests were carried out on pot 1076 #1 and 1069 #1 feeders. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 104 A special single outlet chute was installed to allow greater ability to catch all the alumina from the shot. A large pan on a stick was used to catch all the shot and the samples were weighed. Tests were carried out on a variety of feeders. A series of 20 shots was taken from each standard feeder. Then the feeder was removed and the spool was changed from a standard to an insert design. The feeder was reinstalled in the pot, and a further 20 shots were taken. A summary of results is shown in Table 2-VI and detail comparisons are in Tables 2-VII to 2-VIII. The percentage drop in shot size was similar to that found off-site and to that found in the plant on cold feeders i.e. 23% drop in shot size off-site (Section 2.4.2 and Figure 2- 7) and 26% on-site on cold feeders (Table 2-VI). In almost all cases (8 of 9 runs), the insert gave less shot size variability with an average improvement of 23% shot-to-shot. The S.D. from feeder-to-feeder comparisons was 43% less for the insert. The drop in shot size from "without an insert" to "with an insert" was 23%. Once again, this verifies the insert improves shot size accuracy (despite the fact that the shot size is less). The magnitude (in grams) of the standard deviation was always better with an insert, and the S.D. as a percentage of the shot size was also better. As with on-site cold tests, a S.D. of 3% was found. Operating pot testing is very difficult due to restrictions of pot operation, crane requirements and awkwardness of catching all the shot on an operating pot. 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(0 „ -J 1 UI O CM co co ^r >- z 8 I o £ W 5 i ~ i 11(0 tf) ft 8 e ffi P O Q. i§ D-fe ^ if) E = E c K tig ro .2 .O til ss o ,~. 3-.E c"- S C TJ P 5 cc -CSI< ocsi ^ 1 - I c.3? a g CO £ dra d - 1 TJ T to ^r to to 1* •QT raCM CD j; CM in CM 8 8-2= 8 fe co ? in s 8 Ui T)Js o s &S •a tf) 2 o p •o P 60 > S raP JoS 8 — (0 jr fe ro fe >- V ifsli -, TJ ra TO fe ti in -i III 1 E o _J r- fe © to p J^ ro CJ z O §£Ec»o) r- 2 o Z = S3 o, TJ g y ro ^ ro — QJ •^ -£ < or ^7 < LU 7" h- < UJ T h- IB O-» »~ lzE c5a to JR t5 -Q.„ = 0) UL CJ III (O LuL CJ III Uu<_ UI o (/o) mO) P CD £ P £ TJ CO c E < O Z 0 Z b J; Q (9 tr in tr m CO JJ D)^-.tf2^j-E fe X in l- I i- U) tf) tM fe y Ife. rr t/oJ o LaLc CoO p o _tf) J7) p 0 c/> z o u. to u co N in ^S ra P ro en 3 5 5 5 8) o iS HI < UI o EE o o s 3 3 =- ro or Q or co co UI W O UI J 0. a CO t/j H o CN 04s CO u CHAPTER 2 SHOT SIZE ANALYSIS PAGE 108 In addition, Table 2-VIII shows comparison tests of different dwell times and different speeds. The Is dwell time change affected the standard feeder by 5% versus 1% for the insert. The change is speed affected the standard feeder by 11% but had no significant effect if an insert was installed. All these tests verified conclusions reached for off-site testing. The insert had achieved better performance of shot size accuracy and shot size standard deviation for the worst case testing i.e. on an operating pot. Also, the tests verified that cold testing is satisfactory to give trends and that operating pot testing is not necessary unless one wants to test the complete feed system. Thus, in conclusion, the assumptions made at the start of the study (Section 2.3.4) have proved to be valid. 2.6.4 Portland - A3 Feeders Shot size accuracy for A3 and AEDD feeders in operating pots was conducted to compare their relative accuracies (Table 2-IX). This showed that the standard deviation for the A3 was only 21 % of the AEDD feeder-to-feeder and ws% of shot-to-shot on hot pots. Note that the S.D. for hot pots was several times that of off-site testing, possibly due to hold-up in the chute and inaccuracies catching the shot when anodes were close by. The shot size for the A3 varied little for different dwell times, but the AEDD was much more sensitive. Thus, the independent A3 feeders were more accurate and less variable than the integrated AEDD feeders in operating pots. It is interesting to note that the accuracy of the AEDD with insert, the sequential and A3 feeders is better than the Pechiney and Boyne integrated feeders which are quoted as lkg (+/- 50g) (ref 101) and 1.8kg (+/- lOOg) (ref 102) respectively for non-operating feeders. Thus, the insert in the AEDD has taken a feeder with poor accuracy and raised it of the best in the world. CHAPTER 2 SHOT SIZE ANALYSIS PAGE 109 TABLE 2 - IX A3/AEDD SHOT SIZE COMPARISON (OPERATING TEMPERATURE) Feeder Dwell Mean(g) Feeder-to-Feeder Shot-to-Shot Number Time (s) S-D.(g) S.D.(%) S.D.(g) S.D.(%) of tests AEDD 2.5 1318 55 4.2 88 6.7 15 4.5 1299 112 8.6 64 5.0 15 Ave 1308 89 6.8 77 6.8 A3 3.0 1245 17 1.4 37 3.0 14 4.0 1250 16 1.3 36 2.9 14 5.0 1249 20 1.6 48 3.8 14 Ave 1248 17 1.4 40 3.2 A3/AEDD 19% 21% 83% 48% Note: (i) Samples taken at weekly intervals over 3 weeks to include variability due to alumina size (20 April to May, 1994). (ii) All feeders had single outlet chutes. (iii) All AEDD feeders had Atlas Copco cylinders and inserts. 2.7 EFFECTS OF PNEUMATICS ON SHOT SIZE The pneumatics of the pot can affect shot size accuracy because of the effect on piston speed. The pneumatics of the feeding system includes cylinder size, number of cylinders per DCV, DCV size, air pipe length, air pipe size, exhaust muffler size and condition. For Portland pots, there are 5 feeders operating off the one DCV yet off-site tests used one feeder per DCV (as is the case with most independent feeders in operating plants. In order to quantify the effect of multi-feeder/DCV versus one feeder/DCV, one feeder (Parker 5001 cylinder with #2 Portland assembly) was tested off-site and on-site under cold conditions. The alumina that was sampled in the on-site operating pot was retained and reused on the testrig using , once again, the same feeder. Equivalent stroke times were used to cater for different response times of the valves. The results are seen in Table 2-X. ^^ Z r-Z -J Z Ul < < 9 LU CO O (0 O O CO O CO UJ o o UJ LU LU STI C z >• >- z z z z z >- z >- IGNIFI C AT I t- CO o CO to §. Q CO CO CO CO CO CO u UI *wr o oCN cCoO CO CO CO m CO O co m tf to O # to I S~+*. o 00 CO •>- CO CO •> ^ H o o oo CO •tf co CO o c "c ' LU % CN •tf CN CD c c TJ T- CO co "" >- O O CO 8 8 p w o c CO o OH a. o ,_». - •a T3 > •*««c ! c Ig to CO (0 o 05 Ul ?o o to UJ s ro in to 1- to to to to to to to to to to UJ _hi i- k_ co _J 1 in CJ T-^ o m o m o m o o m < z 2. LL _l CN CN CO •tf in •tf O cd tf co" z o < CO LL. UJ Ul 0. H o to o § OC co LL a. Q Ul O 1- ro II z O LL OC CD c a £ T- CN 2^>p a CO CO orto (0 to to to 5s o UJ $ra h- 05 H Ul to CO co 45 z l2- 1Ul- = r _i to to to to LU o •tf JU O _i CN CO 5 r- to S Z> > ion (Si c 1 > LLi p a o z LL ^-^zf^r o Ul O —A-*' o Both on-site and off-site tests improved with and without inserts. The variability off- site was about 60-69% of on-site testing using exactly the same feeder and the same alumina. It was concluded that the reason for the variability is the different pneumatic effects of the "single feeder/DCV,r off-site versus "multi-feeder/DC V" for on-site tests. Tests on sequential feeders showed that the shot-to-shot S.D. was 0.5% (Table 2-II) versus 1.1% for the same feeder with an insert (Table 2-X). In both cases the same feeder assembly and spool were used, so the difference must be because of the pneumatic control of the feeder. The valve used on the off-site test rig was about 2 metres from the cylinder yet the sequential feeder had the DCV mounted on the feeder itself. Thus, the design of the pneumatics of a feeder has a significant effect on the accuracy of that feeder (Section 4.6). This establishes that there is a major difference between the speeds and reaction of the feeders with valves close and far from the DCV and that one feeder/DCV is superior to multi-feeders/DCV This also illustrates the effect that the proximity of the valve to the feeder has on shot size accuracy. The closer the valve to the feeder, the better the results. In studies of the response time of feeders to DCV activation (Section 4.6), the performance of five feeders in a pot was dramatically different from the situation where there were only one or two feeders in the pot. Stroke speeds are much faster if less than 5 feeders are installed...about 0.5s for two feeders versus 1.6s for five feeders. Hence, the operation of 4 or less feeders in a Portland pot will affect average shot size to the pot. This suggests that "dead" feeders need to be changed as soon as possible after they are detected, rather than just rely on the pot to adjust itself. 2.8 DIRECT FEED INTO THE HOLE The standard AEDD feeder chute is an open design with no direction of flow towards the hole (Figure 1-9). Study on the Alcoa feeder has found that up to 10% of the shot does not reach the liquid bath due to blowing of the alumina away from the hole (27). CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 112 The concentrated stream from a single or double outlet chute minimizes this loss. It is inefficient to optimise the dosing unit to achieve under 3% accuracy if there is up to 10% loss after the dose is metered. As discussed in Section 1.2.13, the preferred design for good dissolution of alumina is fast dosing into the bath. This occurs in independent feeders where the plunger is out of the way of the dose (Figure 1-12 and 1-13). The common independent feeder designs add a horizontal velocity component to the alumina which should help mixing in the bath. It should be noted that despite all the activity on chute design, the hole is often partly or fully closed in most operating pots. T17 estimates few holes are open, Til estimates less than 30% and Portland had 30% of holes open (refer Section 6.3.6:3). D9 (which has AP18 independent feeders) ensures that each hole is opened every 32 hours, but many holes are usually closed. A single outlet chute aimed at the holes directly assists the alumina in the hole area either to (i) fall immediately into the bath (good mixing), or (ii) allow the alumina to be directed into the path of the plunger on the next crust break. The latter results also in good mixing in the bath as the alumina is pre-heated prior to entry into the bath. In fact, at Dl, the feed/crustbreak cycle is set so that the dose occurs immediately prior to crustbreak so the crustbreak actually forces the alumina into the hole. The AEDD feeder tends to spread the alumina around the feeder hole, so alumina cannot feed directly into the hole (Figure 1-10). However, when it does spread around the hole, it is preheated which is good for dissolution when it does slip into the hole. It is also desirable to have the chute outlet as close as possible to the hole to minimize dust loss. However, the closer the chute to the crust, the greater the chance of build-up on the plunger which in turn may foul the chute. Unless the cylinder is strong enough, CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 113 the plunger may not fully retract. For integrated feeders this causes the alumina to leak or not to feed. Pechiney has a "dag scraper" to cut off the dag. Examination of Figure 1-11 and 1-12 shows a shroud around the crustbreaker plunger to clean off dags. However, this is still not effective for new pots where dags often occur (Figure 1-2 and 1-3). Dag cleaners in Alcoa and Alcan smelters have all proved ineffective and most have been cut off. This is due mainly to the small cylinders (100-125mm O.D.) versus the large cylinders (160-200mm O.D.) on independent designs. The latter cylinders have about 4 times the force to knock off the dag. At Portland there is insufficient space between the dosing unit and the plunger hole to fit a single outlet chute with angles over 60 (as suggested by Arnold (44)). A single outlet chute with an angle of about 45 was tested but alumina tended to hold up in the chute (Figure 2-1). Alumina discharge tended to stick on the wet plunger and buildup dags which in turn fouled the chute. About 5 different chute designs have been used at Portland to address optimum dose delivery (Figure 1-22). The final double outlet chute largely solves all the above problems by raising the chute above the high dag level to stop fouling and allow the alumina to pass under the plunger into the hole (Figure 2-17). This chute fits the AEDD, A3, sequential feeder and virtually all AEDD feeders installed at Alcoa/Alcan/Alumax plants. However, comparison of 10 pots with double outlet chutes showed no statistical difference in performance to standard AEDD chutes. It could be that the preheating effect may be more important than the direct feed that Reverdy (24) recommends. If independent feeders break a hole every 2-3 feeds, the alumina tends to bridge the hole (unless there is a large plunger penetration distance). This negates the desired benefit of direct feed suggested by Reverdy (24), so Pechiney feeders have a deep plunger penetration to ensure a hole is continually provided at regular intervals. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 114 However, this contributes to greater plunger wear (refer Section 5.4.2). To correct for this, Pechiney use stainless steel plungers to ensure long plunger life. FIGURE 2-17 A2 FEEDER WITH DOUBLE OUTLET CHUTE CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 115 2.9 FACTORS AFFECTING SHOT SIZE ACCURACY IN POTS Based on the discussion in the previous sections of this Chapter, there are many factors that can affect shot size accuracy in pot alumina feeders; a summary is given in Table 2- XI. Integrated feeders are significantly affected by 32 factors and independent feeders by only 20. Clearly there is much less chance of variability for the latter design. The importance of these factors varies from plant-to-plant...even for the same type of feeder design. For example, the results of Portland were worse than T6 and Til, despite all three being so called "identical" AEDD feeder designs. Even within one plant, the variability from feeder-to-feeder can be very significant. In the case of two Til feeders tested, they had quite different characteristics due to the quality of cushioning and Portland had different results for Terry cylinders and Parker cylinders. Thus, each factor needs to be addressed by each plant to determine if they have a problem. Routine shot size testing needs to be done at all plants; this is rarely undertaken (except for Portland since January 1993). However, for both integrated and independent feeders, significant improvements can be made by simple attention to quality of manufacture, operation and design. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 116 TABLE 2-XI FACTORS THAT AFFECT SHOT SIZE INTEGRATED INDEPENDENT POT - DCV condition - variable supply pressure - Mounting - loose fittings shake feeder - Restricted hoses - Muffler quality - Non vertical mounting - Hold up in chute under super - Restricted flow control valve - Alumina loss between feeder and liquid bath ALUMINA - Angle of repose i.e. 0 - 55 degrees - Angle of discharge i.e. 60 - 90 degrees - Alumina restricted to feeder - entry, kidney plate - Insufficientflow time to fill feeder before next shot - Fixed computer dwell time yet variable feeder response time CYLINDER - Different cylinder designs have different movement characteristics - Speed down over first 50mm and last 50mm of up stroke - Good top cushion causes alumina by-pass, - Seal type .... Teflon allows faster piston stroke speeds than Viton - Age of cylinders ....New ones are slower - Seal condition ...Leaking rod seals cause fast speeds down and slow up - Seals catch and release randomly - Piston not perpendicular can cause vibration - Foreign objects in cushion ports - Tight clearances under heat ASSEMBLY - Spring tension... low tension increases variability - Spool alignment not central - Loose objects jam spool - Non central movement - Alumina jamming spool - Cup jammed on assembly - Tight clearances under heat - Spool bottom face not horizontal causes fouling of spool - Spool not central causes fouling of spool MEASUREMENT - Not collecting full shot - Dust loss - Weigher gauge capability - Not consistent dwell time - Hold up in chute causing carry-over to next shot 32 20 CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 117 2.10 MAJOR FINDINGS FROM SHOT SIZE ANALYSIS (i) The shot size from Portland feeders was affected by alumina particle size (13%), different feeders (13%), type of cylinder (9%), dwell time (7%), and stroke time (2%). Up to 37% difference between one shot and another on different feeders was possible. (ii) Spool inserts reduced shot-to-shot standard deviation from 7-14% to 2% at the original Portland stroke time. (iii) T6 feeders showed little variability due to poor cushioning on the cylinder. (iv) One Til feeder was similar to Portland feeders with Parker cylinders (poor) and one similar to Portland feeders with Terry cylinders (good). T6 operated like a Portland AEDD with Terry cylinders. (v) Spool inserts reduced shot size S.D. (generally by at least 25%) for Portland, T6 and Til feeders. (vi) 53mm inserts had better results than 81mm inserts but reduced the original shot size by about 25% (versus 17%). (vii) Inserts reduced the sensitivity of feeders to spring and cushion quality. (viii) Independent feeders are generally more efficient then integrated feeders for shot size variability. (ix) The AEDD with insert, the sequential feeder and the A3 are more accurate than the Pechiney and Boyne feeders. (x) Seal life and feeder age affect feeder speed and shot size; old feeders are more accurate. (xi) 32 factors may affect shot size accuracy for integrated feeders including mufflers, feeder location in pot, cylinder type, assembly, spring, and mounting. Independent feeders may be affected by 20 factors. (xii) Shot size testing on operating pots found that one feeder/DCV is more accurate than multi-feeders/DCV. (xiii) Shot size testing on operating pots verified the off-site testing. (xiv) Cold new feeders on Portland pots showed 61% improvement in S.D. shot-to- shot and 36% improvement in S.D. feeder-to-feeder if an insert was installed. (xv) Chute design can affect effective shot size into the liquid bath. CHAPTER 2 SHOTS SIZE ANALYSIS PAGE 118 (xvi) Independent feeders and integrated feeders with sequential dosing are superior to the integrated multi-feeders/DCV feed systems with respect to shotsize accuracy and standard deviation. 2.11 RECOMMENDATIONS FROM SHOT SIZE ANALYSIS (i) Spool angles to be 60 at top and bottom to improve shotsize variability. (ii) For multiple feeder operation with one valve, minimum dwell times should be set for new feeders at the farthest position from the control valve such that the movement time is complete to ensure complete shot is delivered. (iii) Ensure that seals or bush clearances are sufficient to not hamper the speed of newly overhauled cylinders. (iv) Install high temperature springs on AEDD designed integrated feeders to ensure quick operation of the spool, to avoid inefficient sealing of the spool and to cope with higher temperatures. (v) Overhaul feeder cylinder cushioning routinely on refurbishment to ensure no effect on piston speed and to minimize fatigue. (vi) Install cushioning testrigs usin g accelerometers to test feeders after overhaul. CHAPTER 3 DOSE DELIVERY PAGE 119 CHAPTER 3 DOSE DELIVERY 3.1 SUMMARY A survey was carried out at Portland to establish the reasons why crustbreak plungers do not break a hole in the crust. The cause of nearly 900 blockages were traced. The major causes were new set carbon (28%) feeder repeaters (24%), low bath (19%), pot repeaters (12%) and pointed plungers (7%). "Repeaters" were feeders or pots where there was a continuing problem of blockages. Feeder repeaters comprised only 2% of total feeders installed. Only 3% of pots were repeaters but these comprised a total of 36% of blockages. New set carbon was also the main cause at other smelter locations which conducted similar surveys. For surveyed pots that had several anode effects per day, about half were due to problems related to blocked feeders. The major cause of these anode effects was turning off the feed and no one turning it back on in time. This resulted in about 20% of the anode effects. Suggestions for improvement included longer dwell time after setting, better anode cover practices and close control of pneumatics quality. Investigations into pressures and kinetic energy of crustbreakers at 13 smelters showed that high kinetic energy is not necessary to ensure the feeder hole is clear. Air cylinders of 125mm O.D. are adequate if attention is taken to cover practices, pressure at the crust face and dwell time, but a 150mm O.D. size is a safer option to cope with pot variability in bath level and anode cover practice. CHAPTER 3 DOSE DELIVERY PAGE 120 3.2 INTRODUCTION A basic requirement of a pot feeder is to ensure that a hole exists in the crust so that the alumina can enter into the liquid bath. Generally, most plants standardise on a minimum of 50mm O.D. plunger width closest to the crust. This basic question has largely been overlooked in the design of point feeders. Close examination of material at the target area shows this is not a "crust". The area around the hole is a hard fused crust, but the hole area is mostly a bed of loose particles of anode cover or alumina in the path of the plunger. If this really was a "crust", then how did it form so quickly... within 3 minutes? Normal hard crust takes many hours to form. For example, the hole which is used to tap metal can be broken by a crow bar even after 24 hours. The duty of the plunger is to penetrate this material in a vertical path. This understanding is critical to design the crustbreaker adequately. High kinetic energy (KE) plungers with pointed tips surfaces are not a good means of penetrating a granular material. Pressure by itself also may not be effective if the material arches significantly. There are several techniques that the industry has used to "guarantee" a hole, but these all have drawbacks (Table 3-1). TABLE 3 -1 TECHNIQUES TO BREAK A HOLE IN THE POT CRUST Principle Method Drawback High pressure large cylinder high air usage, cost small plunger O.D. hole can bridge, short life High KE fast speed fatigue, wear high mass fatigue Deep penetration dip plunger in bath plunger wear CHAPTER 3 DOSE DELIVERY PAGE 121 One solution to all blockages is to increase the O.D. of the cylinder to such a size that the pressure is so large that the plunger will always pierce a hole. Pechiney use large cylinders (200mm O.D.) versus 100-160mm in non-Pechiney smelters. The Alcoa experience is that 100mm is adequate but 125mm gives a little more insurance (as verified by Portland tests). It is desirable to optimise breakage rather than increase cylinder size. The latter is not an option at most plants. This Chapter examines why holes block and a comparison is made of the high pressure and high KE plunger operation. 3.3 IMPACT OF ANODE COVER ON HOLE BLOCKAGE Arnold and Gu (79) examined the material flow properties of anode cover and the factors that could affect the hole blocking. Conclusions from this study included: (i) Cover Size Large particles bridge easier than alumina, but a mixture of both is even worse; even for holes up to 100mm diameter. Anode cover size is selected according to pot design and heat balance considerations. Many smelters use alumina or minus 3mm material as anode cover; this material is unlikely to bridge easily. The minus 20mm sized anode cover used by Alcoa/Alcan/Alumax plants bridges more easily. Generally there is little choice of a plant to change the size of anode cover as the mix is a culmination of a material balance of returning anode cover from old anodes. (ii) Gas Flow Bridging will occur more easily if there is an up flow of gas. The better the overall cover practice on the pot, the faster the gas flow through the feeder hole, and the greater the chance of blockage. A "gas hole" would be beneficial i.e. a hole always placed in the cover to let the gas exit away from the feeders. CHAPTER 3 DOSE DELIVERY PAGE 122 (iii) Cover Height The deeper the cover and the finer the material, the more difficult it is to penetrate the cover. Once again, the more the pot is covered to reduce heat loss, the worse is the effect on feeding. If alumina or fine bath dust is used as cover, it is more difficult to break a hole. (iv) Minimum Plunger Size A 30mm O.D. plunger penetrates the cover more easily than a 50mm O.D. plunger. This observation verifies the reluctance of operators to remove pointed (50mm O.D. or less) plungers. If it still breaks the hole, they prefer to leave the feeder alone. This action has the drawback that the hole may block due to a size change in anode cover or a drop in crust height resulting from a drop in liquid level in the pot. (v) Hole Size The hole size required to stop bridging varied significantly based on cover depth, cover size and gas flow. Virtually any hole will bridge if given the worst of these conditions. Clearly, the smaller the available hole area, the greater the likelihood of bridging. (vi) Vibration Vibration significantly increases the chance of bridging. This is usually not a problem in a pot as there is almost no vibration present. CHAPTER 3 DOSE DELIVERY PAGE 123 (vii) Draw down angle This angle adjacent to the hole was 40-45°(from horizontal) for all types of cover tested viz. <4mm (mainly alumina size), >4mm, and 50% mix. This suggests that the hole will be unlikely to block if the angle of cover is allowed to be at or below 40° so no material can free fall into the hole. Integrated feeders have a crater over the hole. An extreme example is that shown in Figure 1-10. Independent feeders feed directly into the hole so there is little to no crater (Figure 1-13). Thus, independent feeders are less likely to block. It is concluded that the mechanism of blockage is one of flow of alumina and anode cover into the hole and bridging. The better the cover practices (deeper, less gas holes) the greater the chance of blockage. Hence, the most likely times that blockage problems will eventuate are (i) directly after covering anodes when anodes are changed (set), (ii) after pot dressing (general maintenance of cover material over the anodes), (iii) after rapid feed (usually after anode effect), or (iv) if the cover is vibrated by bridge movements (during tapping of metal or resistance change). If a granular mix is used as anode cover there is greaterrisk o f blockage. This theory is verified by the observation at all Alcoa plants that the greatest frequency of blocked feeder holes is directly after the anodes are set and covered. Portland pots have some of the worst conditions for blocked feeder holes viz. large anode cover size, 62% of anodes are set next to feeders (versus 30-50% for most other pots), small gaps between anodes where feeders are located (due to high percentage of carbon area in the pot). A survey of causes of blocked feeder holes at Portland identified the situation when blockages occur; the results of this survey are discussed in Section 3.5. Results have since been proven to be similar at other Alcoa plants. CHAPTER 3 DOSE DELIVERY PAGE 124 3.4 CRUST BREAKAGE TECHNIQUES The Alcoa/Alcan/Alumax technique for crust breakage has generally been by use of high kinetic energy. Hence, high plunger speeds have been a feature of their recent plants except for T4 which chose slow speeds. Plants such as those of Pechiney, concentrate on high pressure (150-200mm OD cylinders versus 100-125mm for Alcoa/Alcan/Alumax) and low velocities (3s stroke time versus 0.4-1.0s respectively). Alcoa/Alcan/Alumax integrated feeder plants not only have smaller cylinder diameters than independent feeder plants, but also often have lower mains air pressures. However, these feeders still break holes successfully. This suggests the cylinder diameter preferred by independent feeder plants is excessive. Use of large diameter cylinders results in high air usage, so these plants needed to develop a feeder to compensate for this...the independent feeder. Which technique is better...high KE or high pressure? Sixteen smelters participated in timing the stroke times of their feeders (75). This was used to calculate kinetic energy, to ascertain if there was a minimum level required for feeders to break a hole. The mains pressure at each plant was combined with the plunger diameter to calculate the pressure at the crust. Results are summarised in Table 3-II and Figure 3-1. CN I—I W ^^ a TJ O UJ (C0 O •e a. OOOOOOOOOCJOOO »- »- 6 0 0 d r— Q. ^ 5c3 in w ra > UJ a. h- fK o a. "TO ooooooooooooo 0 0 0 _. 0 _ 0 LLUJ. 0 LL ja^. 0(0(ON«10(OttlCII)(D(!nnnoionnoioinnnp ) in a> 0 8 a> 9 UJ CM * « > O , , a: =5 a. *^«N;««pCM0q^^''«J a>.CO 0 ID sn 0 v CO UJ c 3 OOOOOOr^^oOOO*- d d ^ d T^ ^ d UzJ r« UJ £ Hl > t. * 0) z E TJ CL 0 c -j 3 qininsffl^spiT-^'tNC) 01 1-; CM *- 00 - !-H >< Q.E ss_l >.2 1 0 O 1- Cl 1- « (9'91)9U <9'91)9U. rt ui 0 w U z a- * O ai e r e Ml "" <»-n)9u <»-n)9u r e z D •£ 111 th e s u s u nt s P w i . o c. —«••• O U ^ :m 0 r < Q u (8'9'9"l)9U - °- ,_-' a°) •» HH in -i (8'9'91)SU *~^ ,_ j; •> (a - _ ~ o> *. H 5?. •2 „ •= = <= U4 0 — u i? (U Q UVnJsu. (z'»-n)su E - o c Z N <» > £ a ° • = Z Pechiney's specification for crustbreak cylinders is 0.5-0.7 m/s between stops. However, most Pechiney smelters operate at much slower stroke speeds. (In Table 3-II, the peak velocity is double the average velocity.) Stroke speeds of this study varied between 0.2 m/s and 0.9 m/s. Note that the slowest rates are for 125mm cylinders (except Portland and T10). If one considers KE values on 125mm O.D. plants alone, Portland could drop KE by a factor of 10 which relates to a stroke time 3.2 times the original stroke time. Thus, a 2.5s downstroke appears acceptable for Portland to break a hole based on data from operating plants. One would surmise that high KE is not the most critical factor in breaking the crust. If it were, then all plants could slow their feeders down to that of the lowest KE plant, and benefit from lower friction wear on components. Generally, wear rate of seals increases by the cube of the velocity, so any piston speed reduction is expected to have significant benefit if friction is the root cause of seal wear. Most plants would be very nervous about reducing cylinder speeds for fear that feeder holes would block and anode effects would increase. 3.4.2 Pressure The pressure applied by most of the integrated feeder plants is similar and is the lowest of the plants surveyed. Using a 125mm (not 100mm) cylinder greatly enhances crust pressure; by 56%. In addition, the plants using 125mm cylinders (and T7) also have CHAPTER 3 DOSE DELIVERY PAGE 128 higher supply air pressure. The combination raises applied pressures to double that in integrated feeder plants which use 100mm cylinders. Once again, there is quite a difference between plant pressures and yet all plants survive. However, the smaller cylinders will be expected to have more problems with crustbreakage from pressure. If increasing air pressure is not possible, then reducing plunger diameter may ensure crust breakage for lower pressures, but at the expense of plunger life. 3.4.3 Stroke Time to Break the Crust In 1991, when Portland was testing slow crustbreak speeds, there were very strong opinions within Alcoa and other companies that this was very dangerous due to the belief that high kinetic energy was needed to break holes. For the previous 30 years, feeder design and operation was based on maximum KE in Alcoa and plants that used Alcoa feeders e.g. Alcan and Alumax. Initial tests were undertaken almost in secret until there was clear evidence that speed was not important. Only when stroke speeds as slow as 8s (versus 0.9s normal stroke) had been proved successful was senior management informed that slow speed was possible. Even then, it was only the fact that test had been successful for 3 months that enabled them to be convinced. It is hard to disagree when the answer is "We've already done it". As an extreme, tests were conducted at zero velocity (Table 3-HI). In the tests, 29 pots had the feeder turned off for periods of 2-7 hours. After the stated period, the plunger was gently lowered to the top of the crust. Then the air was turned on and the time to break through the crust was timed. Instantaneous breakthrough occurred for 90% of feeders after 4 hours. Crustbreakers on most potlines operate about every 3 minutes. It seems likely that there is tremendous over capacity in operating pots. Note that in this test there was no alumina or anode cover placed over the hole during the period. This factor is quite significant as it is postulated that these materials are the CHAPTER 3 DOSE DELIVERY PAGE 129 TABLE 3 - m TIME TO BREAK A HOLE AT ZERO VELOCITY TIME TO DOWN BATH POT FEEDER BREAK DATE POT CaF2 TIME LEVEL TEMP LOCATION CRUST (hrs) (s) 01.06.91 4012 160 927 3 3 0 4014 T. 180 972 3 2 0 4018 170 976 3 3 0 4024 150 940 3 3 1 4032 170 954 3 4 1.5 4033 180 960 3 4 0 02.06.91 4012 150 952 3 7 0 4014 170 955 3 7 1 20-26.06.91 4018 200 962 3 7 0.5 2041 4.9 200 958 1 4 0 2042 4.8 150 966 5 4 0 2043 4.9 160 945 5 4 0 2044 4.7 230 935 5 4 0 4049 5.0 120 985 4 0 2010 4.7 200 993 4 0 3088 5.0 200 1005 4 0 3091 5.0 210 951 4 0 3101 4.9 200 968 4 0 2040 4.7 190 960 4 0 2011 4.9 200 972 4 0 2004 4.9 230 958 4 0 2060 4.8 190 945 4 1 4075 4.9 270 970 5 4 0 4026 4.7 200 994 4 0 4020 4.8 240 969 5 1 2009 4.8 180 961 5 5 0.5 1033 4.9 150 955 5 0 1037 4.9 200 955 5 5 0.5 1034 4.9 200 950 5 0 Notes: (0 Method is to turn feeder and kidney plate off for stated period with plunger above crust. Then slowly rest plunger on crust, put air on at zero velocity and time period to break hole. (ii) The lower the CaF2, the harder the crust. A 4.7-5.0% CaF2 is typical for smelters. CHAPTER 3 DOSE DELIVERY PAGE 130 reason why plungers mainly do not maintain a hole. It was concluded that it is not the hard crust that is the problem with blocked feeders, but compaction of the loose anode cover that falls into the feeder holes. The plunger drives the material into the base of the pot until the anode cover packs into a solid blockage. This test enforces this argument as clearly it takes little effort to break the crust per se. It was concluded therefore, that specification of KE or pressure in themselves are not solutions to the task of breaking a hole successfully in a repeatable fashion. The stroke time (plunger speed) of Portland feeders was reduced late in 1991 to a speed that would largely eliminate fatigue of components viz. a change in stroke time from 0.9s to 1.6s for downstroke (Section 8.6.2). There was no change in anode effects due to blocked feeders. This proved that the above analysis was correct (even for a 70% reduction in kinetic energy). T17 has also slowed down their feeders due to the success of Portland and no effect has been seen in anode effects or blocked feeders. There is no commonality between KE or pressure for all the plants surveyed, so these parameters do not appear to be the main criteria for selecting a feeder design that does not block. Yet, blocked feeder holes are a common problem in all plants. It is postulated that the causes of blockage are not due primarily to the design of the feeder but to other independent factors. The reasons for unsuccessful breakage are more complex, as the reasons are related to things other than hard crust. A blocked feeder survey was conducted at Portland to identify why blocked feeder holes occurred and how to prevent them. CHAPTER 3 DOSE DELIVERY PAGE 131 3.5 BLOCKED FEEDER SURVEY A survey was conducted from mid December 1991 to mid January 1992 to try to establish reasons for blocked feeder holes. Some 1058 blockages were logged, of which for 9% the cause could not be identified. Table 3-IV shows the results for the 91% of blockages for which the cause could be identified. FIGURE 3-IV CAUSES OF BLOCKED FEEDER HOLES AT PORTLAND CAUSE OF BLOCKAGE NUMBER PERCENTAGE NEW SET CARBON 250 28 FEEDER REPEATERS 207 24 LOW BATH 168 19 POT REPEATERS 106 12 POINTED PLUNGERS 60 7 UNKNOWN 41 5 ORE LEAKS 23 3 OTHERS 21 2 TOTAL 876 100 Notes: (i) 44 feeder repeaters (2% of feeders) had 24% of blockages. (ii) 11 pot repeaters (3% of pots) had 12% of blockages. (iii) Observations were taken in December 1991 and January 1992. 3.5.1 Examination of Blocked Feeder Survey Data Before the survey was conducted, several operators said it was a waste of time. They already knew the main reason was "new set carbon"; that is, the holes block shortly after the anodes are set and covered. While this proved to be the case, it was generally not predicted by most engineering and management personnel across many smelters. Most engineers thought the main reasons would be pointed plungers (where the diameter is small) or when bath level is low (when the plunger does not reach the crust). All nine (9) Alcoa plants have since done similar surveys to Portland and have also found the major cause was "new set carbon". Pointed plungers and low bath level were well down the importance list. CHAPTER 3 DOSE DELIVERY PAGE 132 This is an important anecdote as it illustrates how important it is for engineers to consult operators. Operators have a feel for problems, even if definitive data does not exist. Examination of the blockages shows some feeders blocked frequently; whatever the reason. It seems reasonable to group these blockages as being due to a faulty feeder. "Feeder repeaters" were defined as where a feeder had 3 or more blockages during the survey period. Also, groups of feeders in a pot had frequent blockages. "Pot repeaters" were defined as where 3 or more feeders had 2 or more blockages during the survey period. The prioritising of blockages were (in order with most important first); (i) feeder repeaters, (ii) pot repeaters, then (iii) the stated cause viz. "new set carbon", "low bath", "pointed", "leaks" and "others". Any blockage that had none of these was listed as "unknown". Below is a discussion of the causes and proposed actions for each of the common causes of blocked feeder holes. 3.5.2 New Set Carbon When the anodes are set and covered, the plunger is actually below the top of the carbon and is generally buried by anode cover. At most, there may be 150mm of space between the bottom of the plunger and the top of the anode cover. Hence, the plunger hardly starts moving before it hits the anode cover. So, speed is not the major cause of blockages after setting. CHAPTER 3 DOSE DELIVERY PAGE 133 Anode cover pours onto the anode and flows down into the open hole. If one covers immediately after setting, the anode cover fills the hole from the bottom of the pot to above the anode. When the plunger is extended, it compresses the material as far as it can, but generally there is little movement. It is only when the metal/bath washes the material away at the bottom does the plunger have any chance of "extruding" the anode cover through the hole. In the meantime, every 3 minutes or so, another kilogram of alumina is added to the hole area as there is nothing to tell the feeder to stop operating. Hence, the hole starts to bridge up, being assisted by the compression action of the plunger each crustbreak cycle. In a couple of hours, the pile is up to the chute and there is virtually no plunger movement. From mid 1993, Portland changed from covering anodes immediately after set, to covering 4 hours after set. Blocked feeder holes dropped by 85% after this change. Covering later lets the hole crust up a little so that the cover material only fills up above the crust. Hence, the next crustbreak has somewhere to push the cover material into the pot and the hole has a greater chance to remain open. Note that this technique also reduces the amount of unmetered alumina dissolving in the bath and reduces the amount of bath generation from the anode cover. In the Portland survey, the worst blocked feeder frequency was where operators covered the anodes with the highest anode cover depth. This is a potential problem if it is decided to increase the amount of anode cover to minimise airburn on anodes. T7 found 40% of blocked feeders to be due to newly set carbon in 1992. Now they have the people who set and cover the anodes, ensure that the feeder hole is clear before they leave the pot. They have more than halved blocked feeders since this was done. This appears to be the cheapest and most effective solution to this problem. D2 dig out the hole after covering. (D2 have also achieved 0.03 anode effects per day which is the benchmark for anode effect rate.) CHAPTER 3 DOSE DELIVERY PAGE 134 Increasing the dwell time after set (say for 2 hours) will allow the crustbreaker to reduce the angle of the anode cover to its natural angle of repose, then holes will not block as often. Since mid 1993, this practice has been used at Portland with reduced anode effects resulting. A similar program could be used whenever a feeder blocks i.e. increase dwell time for a few hours to ensure the hole is penetrated. This feature is best installed on plants that have the ability for dwell time to be controlled separately on each pot. Older plants do not have this ability. When a blockage occurs, it is wise to stop feeding until the hole is clear. For integrated feeders this is done by shutting off the alumina valve (the kidney plate at Portland). A number of anode effects are due to people forgetting to open the kidney plate after the hole is clear. The computer will automatically increase frequency of feed to compensate for insufficient feed, however, anode effects occur until the pot feed rate matches demand. T17 has seen the situation where the kidney plate was shut for a month before it was detected! The pot was operating satisfactorily, but the frequency of feeding was detected as being very biqh. Whenever anodes are set next to a feeder there is more risk that the new anode cover will block the hole. Hence, the pattern of setting anodes in the pot and the number of feeders installed in a pot will have a large bearing on blocked feeders. It is preferable to have few newly set anodes next to feeders. Every time a new anode is set at Portland it is next to a feeder. In fact, 2 of the 5 feeders have recently set anodes on each side of the feeder at the same time. This is most unusual in prebake pots, and may be the only pot design that has this feature. Some smelters set one anode at a time in almost a random pattern to avoid this problem e.g. Pechiney. Choice of setting style can assist the prevention of blocked feeder holes. (In addition to the effects of the setting pattern used at Portland, Portland also has the highest number of feeders of any pot in the world. As such, it has greaterrisk o f blockages.) Reducing carbon height or reducing carbon area near the feeder has been used by several plants to minimise bridging across the hole e.g. T6, T7. This has proved very CHAPTER 3 DOSE DELIVERY PAGE 135 successful. The carbon horizontal area around the feeder hole often airburns anyway and is a loss of carbon (which costs about $250/t). Thus, not having the carbon there in the first place will save blocked feeders and cost of carbon production. The latter cost can be significant if the plant is making 180,000 anodes per year as is the case for Portland. Thus, there are several methods that can reduce the frequency of blocked feeder holes: (i) longer dwell time (ii) ensure hole is clear after covering (iii) stop feeding after blockage (iv) different setting pattern for new carbon (v) reduce carbon height (vi) reduce carbon area around the feeder hole. 3.5.3 Repeaters Some 2% of the feeders surveyed had 28% of blockages and 3% of pots had an additional 12% of blockages. Some 40% of blockages occurred on less than 3% of pots. From the previous discussion, maybe Portland was operating at high speed to correct for the inadequacies of only 2% of feeders, and wearing out the other 98% as a result. Mechanical factors ("feeder repeaters", "pot repeaters", "pointed" and "leaks") comprised 46% of blockages. The three basic flaws of repeaters are dwell-time time out, lack of supply pressure and back-pressure. 3.5.3:1 Dwell Time Time-out Both "feeder" and "pot" repeaters can be related to too low a speed causing the dwell time to run out before pressure is applied. This is not lack of kinetic energy, only lack of time. Although all pots have the same computer dwell time, the individual pot may be affected by DCV integrity, local low pressure points in the air lines, restricted hoses, CHAPTERS DOSE DELIVERY PAGE 136 tight rod/piston seals, hose leaks, air leaks in fittings, or alumina in the cylinder. All these factors affect variability of speed and hence the effective time that the cylinder is fully extended. Operating at an inflated dwell-time may achieve full shot size, but results in low plunger life and "dags" and doesn't address the cause of the problem. Chapter 4 addresses these factors in more detail. The solution to "feeder repeaters" is to change out feeders that frequently block - whatever the reason. The repair people can trace the cause and report back to the initiator via a tracking system for follow-up. 3.5.3:2 Lack of Supply Pressure It is useful to monitor air leakage rate so that rod and piston seal leakage can be detected before blockages or ore leaks occur. The most important air leaks to address in order to stop blocked feeder holes are when the crustbreaker is in the down position. Air leak checks are often done when the plunger is up, but this does not detect some piston seal leaks or leaks on some fittings. Pistons often have seals pointing in opposite directions and it is indeed possible for a cylinder to leak in one direction only. One needs to check air leaks on the downstroke if one wants to see the effect on crust breakage. Section 9.4 discusses in depth air leakage detection and the costs associated with leaks. 3.5.3:3 Back Pressure Another major factor causing lack of pressure is blocking of mufflers on the exhaust of the DCV. These block due to grease and deposits from the cylinders and air lines. As they block, they cause a back pressure on the cylinders, slow them down and reduce the pressure on the surface of the crust. Portland had a muffler life of 6 months with nearly all these being badly blocked to the extent that the plunger would not go back up into CHAPTER 3 DOSE DELIVERY PAGE 137 position. Hence, the 125mm cylinder could not even lift a 35 kg mass! There would be virtually no pressure available on the down stroke. This problem is more serious at plants which use a common muffler for several cylinders than designs where there is one feeder per muffler. Til found that blocked mufflers halved the pressure at the crustbreak cylinder. At Portland, the back pressure can get so large for blocked mufflers that the muffler can explode (Figure 3-2(a)). To achieve no back pressure, one needs to: (i) direct the exhaust to another place that does not need a muffler, or (ii) use a "straight-through" muffler which does not block, or (iii) reduce the air flow rate by having several small outlets. It is necessary to not only solve the blocking problem, but also to reduce noise levels. There were several possible solutions to this problem: (i) Redirect flow Some 68 pots at Portland had the exhaust directed to the pot fume duct as an alternative to a muffler. This design is costly (about $300/pot) and although able to achieve less than 85dBA, the noise level is a nuisance to people working in the area as the exhaust tends to echo in the building. Some plants (T8, TIO) direct the exhaust into the pot cavity which is a very good idea as there is no need to muffle the exhaust which may affect back pressure on the feeder. (ii) Straight-through or non-blocking mufflers At least a dozen different muffler types have been trailed as part of this investigation ranging from customised units to car mufflers. All have failed due to noise or blockage. Dr. Bob Randall of Unisearch (University of New South Wales) designed a straight-through silencer similar to arifle silence r (Figure 3- 2(b)). It achieves under 85dBA and does not block. These were installed on all CHAPTER 3 DOSE DELIVERY PAGE 138 pots at Portland in mid 1993 and none have failed in over 2 years versus an average life of 6 months for the original Allied Wotan fibre unit. Costs are about $200 each. FIGURE 3 - 2 PORTLAND FEEDER MUFFLERS (a) Allied Wotan M12 exploded and new (b) Portland silencer CHAPTER 3 DOSE DELIVERY PAGE 139 (iii) Reduced air flow This is the method used on independent feeders. Each feeder is controlled by its own valve mounted on, or near, the cylinder with its own exhaust. The reduced air flow allows the use of small off-the-shelf silencers with little back-pressure, but the installation cost adds up for the large number of feeders on a modern plant. 3.5.4 Low bath This appears to be an obvious cause of blockages. The solution also appears simple - don't over-tap bath or metal. Plants such as T17 and D13 do not have the plungers touch the liquid yet still successfully break holes (refer Section 5.4.2). There is no reason to get the plunger wet if one can predict the location of the crust. Operating to a total liquid level is desired, rather than work to a fixed bath level. Total liquid level control ensures a relatively fixed distance from plunger to liquid and crust. Fixed bath level alone does not guarantee this relative distance as the metal level may change from pot to pot or on changing the metal depth targets of the pot. 3.5.5 Pointed Plungers Note that pointed plungers were only responsible for 7% of blockages in Table 3-IV. Despite all the activity and discussion on "pointed feeders", this is not a major contributor to blockages. Operators often suggest that pointed feeders seldom block. A pointed plunger (50mm diameter) concentrates the pressure to 4 times that of a 100mm plunger. As long as the crust is high enough, a pointed plunger will work. Portland installed 30 Inconel 60mm plungers in 1991. After a period of 6 months, the Inconel pipe had corroded away and the plungers were operating on the 50mm steel vertical plunger shaft quite successfully. Hence, 50mm is sufficient for alumina flow as long as cover practice is adequate. Clearly, the wider the plunger the better it is to break a hole and allow alumina entry (Section 3.3(d)). CHAPTER 3 DOSE DELIVERY PAGE 140 3.5.6 Ore Leaks Ore leaks on integrated feeders that block holes are generally caused by very bad air leaks from the air cylinder that aerates the alumina in the feeder. This floods the pot with alumina; generally from the kidney plate. Providing better rod seals and/or allowing a better air passage for venting air leaks are possible solutions (Section 8.4). 3.5.7 Feeder Location Feeder repeaters were less common on #5 feeder at Portland. This may be due to less cover near the tap hole. Pot repeaters had no blockages at #5. This is interesting as # 5 feeder is farthest from the DCV and could have been affected by slower speeds. So it appears that the cover is more important than the manifold layout (Figure 1-21). 3.5.8 Minimum Dwell Time Determination Dwell-time is chosen to achieve two aims: (I) to release all alumina from the feeder (which does not apply for independent feeders), and (ii) to break a hole. Shot size testing has been undertaken on Portland, T6 and Til feeders over a range of conditions (Chapter 2). For the feeders tested, shot size was consistent if the dwell-time was greater than the movement time for full stroke. Tests were conducted on the pressure profile of the inlet and outlet air lines to the feeders at Portland (Chapter 4). They showed that full pressure was not achieved until well after the plunger reached the end of the stroke. These factors need to be understood for each plant to optimise dwell-time. If one chooses too short a dwell time, there is insufficient time to reach maximum pressure to break the crust and all the shot may not be delivered. One gets no feed at all or not enough. Anode effects will result for both situations. Independent feeders can reduce wet time to about 0.2s because they are not affected by the time to discharge the alumina. This compares with 0.5-1.5s for most integrated feeder potlines, and up to 4s in some potlines. CHAPTER 3 DOSE DELIVERY PAGE 141 3.6 RESULTS OF IMPROVEMENTS TO BLOCKAGE PREVENTION From mid 1992, considerable action was taken at Portland to address blocked feeder holes based on the findings of this survey. Covering was changed from immediately after set to the shift after set to reduce the amount of cover in the holes. Dwell time was increased to 5s for 2 hours after covering to accommodate slump of anode cover in the holes. These changes resulted in reduction of blocked holes after set by about 80%. Air leak surveys identified hardware problems which could have contributed to lack of pressure and variability in shotsize accuracy - leaking feeder valves, hoses, fittings and by-passing pot valves being common. By 1993, air leak rate per pot was under 301/minute at Portland which is the lowest in the Alcoa system and possibly a benchmark level. By November 1991, all feeders at Portland had their KE reduced by 70%. There was no change in anode effects due to blockages. This verified that high KE is not necessary to break a hole and suggested that the large cylinders used by some independent feeder plants (e.g. Pechiney) are unnecessary. Smaller cylinders would be satisfactory, with significant savings in capital and air usage. Although 125mm O.D. cylinders are adequate, they may be a little small to guarantee a hole under all plant conditions. Instead of the 200mm O.D. size used by Pechiney, a 150mm size is recommended for new plants. It achieves its task and is considerably cheaper. The reduction in blocked feeders was a major contributor to halving anode effects at Portland from mid 1991 to mid 1993. Although initially it was thought that holes were broken only by high kinetic energy, it is now widely accepted that this is not the case. Thus, this research has completely changed a theory held by many smelters for over 30 years. CHAPTER 3 DOSE DELIVERY PAGE 142 3.7 MAJOR FINDINGS FROM DOSE DELIVERY INVESTIGATION (i) Holes generally block from anode cover bridging the hole. (ii) Bridging is aggravated by increased granule size, cover height, hole size and vibration of anode cover. (iii) Blockage frequency does not relate directly to the kinetic energy or pressure from the feeder crustbreaker but to a range of factors related to pot operation (such as anode cover practices). (iv) At Portland, newly set carbon was the major factor causing blocked feeders (28%), followed by feeder repeaters (24%), low bath (19%), pot repeaters (12%), and pointed plungers (7%). Similar trends were seen at other plants. Repeaters were due to expired dwell time before full pressure is achieved, lack of supply pressure, and/or back pressure. (v) Mufflers blocking cause back pressure on feeders, reduce effective pressure at the crust, and DCV failures. (vi) A 125mm O.D. cylinder is adequate for crustbreak cylinders, but 150mm O.D. is the recommended size for new plants. 3.8 RECOMMENDATIONS FROM DOSE DELIVERY INVESTIGATION (i) Dig out the hole after covering new set carbon or after dressing the pot. (ii) Provide a long dwell time after set and after blockages. (iii) Consider alternative setting patterns to minimise the number of new anodes next to a feeder. (iv) Log all blockages with reason for occurrence to trace patterns of failure. (v) Remove feeder repeaters if no other cause is found. (vi) Regularly monitor air leakage rate for each pot to identify feeders with piston leaks on the downstroke. (vii) Install silencers that do not block or redirect exhaust into pot exhaust duct. (viii) Control total liquid level (rather than bath level) to ensure reasonable penetration depth through the crust (not necessarily to get the plunger wet). (ix) Install 150mm O.D. cylinders on new crustbreaker units. CHAPTER 4 PNEUMATICS PAGE 143 CHAPTER 4 PNEUMATICS 4.1 SUMMARY Testing was carried out on the pneumatics of a Portland pot for a range of feeder age, speeds, location of feeder in the superstructure and type of muffler. In all, 144 combinations were tested to see the response time before the feeder moved, stroke time to reach full supply pressure, and time to reach maximum pressure (both for downstroke and upstroke). Calculations were then carried out on the time the plunger is potentially in the bath ("wet") to understand better the causes of erratic plunger wear. It was found that new feeders operated slower than old feeders. Paper type mufflers affected the time it takes to reach full pressure but had little effect at reduced speeds. For a multi-feeders/DCV pot, the feeder closest to the pot DCV initially moved fastest, but then the others caught up by the time the feeders reached the crust (except that a new feeder always was the slowest). Variable wet time was traced to an interaction between mufflers and feeder age. Waiting time on the up stroke was the major factor on wet time for the plunger. The transient stalling of the downward movement just after the feeder starts to move indicated insufficient air supply and identified a major contributor to shot size variability for Portland. Having a DCV for each feeder significantly reduces reaction time of feeders so that wet time can be reduced by suitable selection of crustbreak dwell time. Location of the DCVs close to the feeder(s) reduces air consumption significantly. CHAPTER 4 PNEUMATICS PAGE 144 4.2 INTRODUCTION Shot size testing showed that the speed of the feeder affected the size and variability of the shot (Chapter 2). The dose delivery investigation showed that 40% of blockages are related to repeating blockages on the same feeders or pots (Chapter 3). As will be seen in Chapter 5, plunger wear is related to the time that plungers are in the liquid. Clearly, feeder pneumatics are a major contributor to shot size accuracy and feeder failure. Air usage for feeders is usually the highest single consumer of air in a smelter. At Portland, the feeders account for 30% of the plant air usage which is not unusual for smelters. One can see that attention to air usage is a major factor in the efficiency of not only the pot feeders, but also the plant as a whole. To gain a better understanding of the dynamics of the feeding system, an investigation was held into the pressure trace of inlet and exhaust air to feeders at Portland (Figure 4- 1). Factors considered were: (i) different feeders (runs of 2 new and 2 old Terry feeders), (ii) location of the feeder in the pot (tap and duct end), (iii) paper muffler design and degree of blockage (Allied Wotan Ml 2 muffler and Portland Aluminium silencers, Figure 3-2), (iv) speed of feeders, and (v) multi-feeders/DCV versus single-feeder/DCV. Two pressure transducers were used to measure pressures in the air line of off-line pots at several locations. Movement of the piston rod was sensed by a special test unit mounted on the top of the cylinder (Figure 4-2). A piston was pressurised by a 420kPa chamber that forced a small piston to follow the large crustbreaking cylinder piston. The test unit had a piston rod travelling through the top flange of the crust breaking cylinder and pressing against the cushion spear of the piston. Limit switches were mounted to sense end of stroke by use of an extension of the piston rod above the test unit. CHAPTER 4 PNEUMATICS PAGE 145 FIGURE 4 -1 POT PRESSURE SENSING POINTS PRESSURE TAKE OFF POINTS WIDE OR TAP AISLE RESTRICTOR. (12mm) 1Z^ L A NOTES: BALL (1) In closed position, ball valve ^ly' VALVE has a 12 mm hole. NARROW (2) Line A supplies air to force AISLE plunger up. (3) Line B supplies air to force plunger down. RESTRICTORS (12mm) GATE VALVE IR SERVICE UNIT -BALL VALVE (BLEEDS TO ATMOSPHERE) -FILTER -LUBRICATOR CHAPTER 4 PNEUMATICS PAGE 146 FIGURE 4 - 2 STROKE TESTING UNIT <=0 TOP LIMIT SWITCH BOTTOM LIMIT SWITCH EXTENDED PISTON ROD CANISTER TRACKING ROD FEEDER CYLINDER CRUST BREAKER PISTON CHAPTER 4 PNEUMATICS PAGE 147 The piston diameter of the test unit was 32mm. This relates to a 4% change in force on the crustbreaking piston. It was not considered that the unit significantly affected the performance of the crustbreaking cylinder and, as such, could be used to sense the tatter's performance. The pressure transducers were wired to a high resolution chart recorder and traces were obtained of the response; Figure 4-3 shows typical curves. An off-line pot was selected (#1015) for most of the final testing after experimentation on several pots showed that the location of the pot in the rooms was not significant i.e. a feeder in one pot had a similar profile to that in another pot. Two old and two newly overhauled feeders were tested in location 1 and 5. Supply pressure was 800+/-20kPa at 140kPa per division on the vertical axis of the pressure traces. Unfortunately, the figures do not always show each division clearly due to the quality of the copying of the instrument graphs. 4.3 GENERAL OBSERVATIONS ON PRESSURE TRACES Looking at the traces in Figure 4-3, one will see locations noted as location A-L. The cycle occurs as follows: A. Start of cycle: air was applied to top of cylinder. B. Start of cycle: air at bottom of cylinder was opened to vent line of the DCV. A-C. "Waiting time". Pressure rose in the supply line to a stage that it was larger than the exhaust air (B-C) until point C is reached. This usually took 0.8-0.9s for restrictors installed (in) or 0.4s if restrictors are not installed (out). C. Pressure across the piston exceeded the friction (called "sticksion") of the seals and bush holding the piston stationary. Sticksion was about 40kpa. The piston starts to move. There was a rapid drop in curve AC due tofilling th e cylinder up with air as the piston starts to rapidly free fall. Then the pressure difference between the two traces dropped to be too small to overcome the friction and the CHAPTER 4 PNEUMATICS PAGE 148 OS u Q rtt'-- w !_.-.;..:« o ti Z CH CH -j- > o —/%-";! o CH u H ti CH LU CH ti tf 0- J.»tl- r»_- C3- —-:i. 5 co UJ Ul NOISIAia/«d>(0»l NOISIAia/w'MOtrl NOISlAt07!*dO? NOISIAia/nOQZ CHAPTER 4 PNEUMATICS PAGE 149 piston slows. Hence a bouncing effect occurred for a cycle or two until the piston starts to move smoothly. Note that this double movement may be seen in the movement of the plunger on slower moving feeders in an operating pot. This "stalling" of the feeders occurred on all tests (even at the original design speed) and suggested that the air supply was not sufficient enough to maintain full pressure at the piston. A larger air manifold across the top of the pot is needed, or the DCV needs to be positioned closer to the feeder. It was less noticeable for new feeders, slow speed (Figure 4-4), and off-site (Figure 4-5). Also at C, one can see the step change in the second line from the top of the graph. This is the trace of the limit switch for top end sensing. C-D. The time that the piston moved the length (stroke) of the cylinder. The pressure trace was almost horizontal. A-D. "Hit crust" time i.e. the time from start of cycle to end of movement. E. Piston reached full travel (as evidenced by the third top trace from the limit switch). The pressure increased at this point. The rate of build up depends on back pressure and type of feeder. Sometimes D & E occur together (Figure 4-3, 4-4). This is because this feeder was the slowest in the pot. E-F. Pressure rapidly rose and fell at point E when the throttling effect of the muffler and the orifices does not apply. At F the maximum supply pressure was achieved. A-F. Time to reach "maximum supply" air pressure. G. "Maximum force" at the crust. Supply air was at maximum and exhaust air was at zero. F-G. The time between "maximum supply" and "maximum force" on the crust took 0.4-0.6s in most traces. A-G. Time to reach "maximum force" since the energising of the DCV. CHAPTER 4 PNEUMATICS PAGE 150 s o -*- .' —". ., -I (0 •11 1 -LJ LifJ—~rl CH coiai«»0 ti z o triifc-'il »* r tin »« r O Tl Tf' T 0 or J_JS»C^ 0 :::±_ |.D o ,, . !«v M..-S o u • 4-IL • • •jr. Tim Q - — * (fl te~;1" I IS ti "IJLI: *1 ° ti >j 2 OM CH JL B i - _ L. —1— _ , CH ti -*• "S±" "~S ** Ht"::-:ta Jl x.-± S U '>. t - 11JL -rH, _J. J_U T IT ti ti NL " tf d ir "i/l^C- .- ?P * ~" y i v' J CH at J- .:a.Lwi-ur ' f*" -J 355^ CH . _2 ir~ ti ~T" . . ! tf 1 I • -H-HI ™J~ • j__r" 3 "'I."it*. X I :±: J- i J-.J zi-it •" Of f*H- UJ •*••-i ii- ._ s or UJ •g :_:±:szxi^ LU -I ••••» UJ \ Y "m 9 = J--3*-:* L I LL S ±Kfcj±tfi rf-i. §§ • •4-4 • - sizfe*- ; LU LU IS .dh NOISIAia/w*10»l NOISIAia/™»10t>l NOISIAtd/pdOZ NOlSIAia/!*dQZ CHAPTER 4 PNEUMATICS PAGE 151 ;AT MJ H CH ti H ti jq:^il H J CH •IT""iIS j i - • ti 5- ti •• < «8 1 i i o ti H / •z i ; H ' 1 12 NH O i ! tn CH i i i \*\ ts. i •>- ; ..... SUJ ! 1 j i •atn z r i , * . i_u. i j •- o r "^LIZ -L"kr CH _ . •F> ti • > J-l/ i •' : i u j. !\L__/a.j - i ;-«•/ \ H LU- f i CtiH i.i •H ,• CH ; ! -' ti .U.Ji •!• L i- tf o • i ti or LU I I J.-J lr I- * ' T ILi.4—: co •. \r LU 1 h- • N: II. LU :! •*.' 0. : s UJ CBT\ UI ;^A—- h- .— NOlSIAia/^>IOfl 05 ••- NOISIAia/!"iOZ u. NOis.Aiayrtwt'i LL. o NOISIAia/lsdQZ CHAPTER 4 PNEUMATICS PAGE 152 D-G. Time that the plunger was potentially in the bath for the downstroke; the "down wet" time. G-H. This time changed from curve to curve depending on when the valve was reversed. Care was taken that the traces came to steady state before the feeder was activated to go upwards. H. Start of upstroke. I. Pressure traces crossed over with a bounce in the supply pressure. This was not so pronounced as for point C due to the retarding effect of gravity on the feeder mass. J. Piston started to move. The up cycle was similar to the down cycle except there was less "stalling" when the feeder started to move. The piston did not start to move until possibly half way along the I-K stretch. The pressure required to start the piston moving was about three times that needed for the down stroke i.e. 140 kPa. J-K. Upstroke movement time. L. Maximum supply pressure upwards. M. Maximum force upwards. The "total wet time" is calculated as the end of downward movement to the start of upward movement i.e. "down wet" plus "wait time" on the up stroke. This is the total time that the plunger is stationary and it is potentially in the bath (wet). Note from Table 4-1 and 4-II that the down wet time was 40-70% of thetotal we t time. People generally think of dwell time as the main factor affecting the time that the plunger is wet, but seldom consider the response time of the plunger when it is retreating from the fully extended position. One needs a quick response time of the piston on the upstroke in addition to the conventional interest in dwell time. In addition to the times listed, the plunger wet time will be affected by the penetration in the bath. If there is no penetration, "wet time" is a misnomer and the plunger is not affected. Most plants have penetration, and the degree of penetration will affect the CHAPTER 4 PNEUMATICS PAGE 153 time in the liquid in addition to the times listed here where times are specified for change of direction only. Refer to Section 5.4.2 on discussion on plunger penetration. Testing was done using Terry cylinders so that the type of cylinder was not another variable. However, as the Parker cylinders now have Terry rod bushes and rod seals, it is expected that a similar performance would apply. There was little to no difference in the performance of any similar pair of feeders. The values of times in Tables 4-1 and 4- II are the average of a pair of similar feeders. Note that at X and Y, the gap suddenly became larger on several pressure traces. This is believed to be due to other cylinder(s) in the group offive reaching the end of travel before the rest. Once the fastest cylinder piston reached the end of travel, the rise in air pressure in the manifold caused the remaining cylinders to develop a greater pressure. This phenomenon illustrates that the cylinders were being partly starved of air; not "pressure limiting" but "flow limiting" which also shows that one cylinder can affect another. This is a problem for pots using multifeeders per DCV. 4.4 METHOD USED TO SLOW FEEDERS When the decision was made to slow the feeders at Portland in 1991 to reduce fatigue. Three alternative designs were tested and rejected. (i) Flow control valves - block with time due to annular seat and small aperture diameter. - poor experience in Alcoa plants. - anybody can adjust them and no one knows when or how much. Uncontrollable. - expensive as needed on all 2,200 feeders ($300,000) (ii) Pressure regulators - anybody can adjust. Not controllable. - iffix adjustment , cannot easily over ride all pots on major plant upset e.g. power off for extended period. - expensive as need two per pot ($150,000) (iii) Reduce air line size - could not go back if unsuccessful. - expensive ($200,000) CHAPTER 4 PNEUMATICS PAGE 154 TABLE 4 -1 EFFECT OF MUFFLER & SPEED CHANGE - LOCATION 1 TRACE NEW FEEDER OLD FEEDER DIRECTION SPEED TIME SECTOR PA NEW OLD PA NEW OLD DOWN ORIGINAL WAIT AC 0.4 0.4 0.4 0.4 0.4 0.4 MOVE CD 0.9 0.8 0.9 0.6 0.6 0.8 HIT CRUST AD 1.2 1.2 1.3 1.0 1.0 1.2 MAX SUPPLY AF 2.0 2.0 2.1 2.0 2.0 2.1 MAX FORCE AG 2.4 2.4 2.8 2.4 2.6 3.4 DOWN WET (2-AD) 0.8 0.9 0.8 1.0 1.0 0.8 UP ORIGINAL WAIT HJ 0.5 0.6 0.6 0.5 0.5 0.6 MOVE JK 1.0 1.1 1.0 0.9 1.0 1.4 HIT TOP HK 1.5 1.6 1.7 1.4 1.5 2.0 MAX SUPPLY HL 2.2 2.3 2.4 2.2 2.4 2.5 MAX FORCE HM 2.2 2.3 2.9 2.2 2.4 3.5 TOTAL WET TIM LOC1 (2-AD)+HJ 1.3 1.4 1.4 1.5 1.5 1.4 DOWN SLOW WAIT AC 0.8 0.9 0.9 0.9 0.8 0.9 MOVE CD 1.9 1.8 1.9 1.1 1.1 1.2 HIT CRUST AD 2.7 2.8 2.8 1.9 2.0 2.0 MAX SUPPLY AF 4.7 4.6 4.7 4.7 4.9 4.9 MAX FORCE AG 5.5 5.4 5.5 5.4 5.2 5.6 (4-AD) 1.4 1.2 1.3 2.1 2.1 2.0 UP SLOW WAIT HJ 1.3 1.2 1.3 1.5 1.6 1.6 MOVE JK 1.5 1.6 1.6 2.2 2.1 2.1 HIT TOP HK 4.0 4.0 4.1 3.7 3.7 3.6 MAX SUPPLY HL 5.1 5.0 5.2 5.1 5.1 5.1 MAX FORCE HM 6.0 5.7 6.1 5.8 5.6 5.9 TOTAL WET TIM LOC1 (4-AD)+HJ 2.6 2.4 2.5 3.6 3.6 3.6 Note: (i) All figures in cells are seconds (s). (ii) Times are the average of 2 old and 2 new feeders respectively. (iii) "New" and "old" mufflers referto new and blocked Allied Wotan mufflers respectively. (iv) "PA" muffler refers to Portland Aluminium straight through silencer. (v)" TOTAL WET TIME" is the time the plunger is fully extended. During this time there is the greatest risk of plunger corrosion. It is calculated as follows: TOTAL WET TIME = dwell time - hit crust time + wait time (vi) "Trace sector" is the time between points in the pressure traces Figures 4-3 and 4-4. CHAPTER 4 PNEUMATICS PAGE 155 TABLE 4 - II EFFECT OF MUFFLER & SPEED CHANGE - LOCATION 5 TRACE NEW FEEDER OLD FEEDER DIRECTION SPEED TIME SECTOR PA NEW OLD PA NEW OLD DOWN ORIGINAL WAIT AC 0.4 0.4 0.4 0.4 0.4 0.4 MOVE CD 0.8 1.0 1.0 0.9 0.9 1.0 HIT CRUST AD 1.2 1.4 1.4 1.5 1.3 1.4 MAX SUPPLY AF 2.0 1.9 2.0 2.0 2.0 2.1 MAX FORCE AG 2.4 2.4 3.2 2.4 2.5 3.2 DOWN WET (2-AD) 0.8 0.7 0.6 0.6 0.8 0.7 UP ORIGINAL WAIT HJ 0.6 0.6 0.9 0.5 0.5 0.6 MOVE JK 1.1 1.2 1.2 0.9 1.0 1.3 HIT TOP HK 1.8 1.7 2.1 1.4 1.5 1.9 MAX SUPPLY HL 2.3 2.3 2.5 2.2 2.4 2.6 MAX FORCE HM 2.3 2.3 3.3 2.2 2.4 3.3 TOTAL WET LOC5 (2-AD)+HJ 1.4 1.2 1.5 1.0 1.3 1.3 DOWN SLOW WAIT AC 0.9 0.9 0.9 0.8 0.8 0.8 MOVE CD 2.0 2.0 2.0 1.2 1.2 1.3 HIT CRUST AD 2.9 2.9 2.9 2.0 2.0 2.1 MAX SUPPLY AF 4.7 4.4 4.6 5.0 4.9 5.0 MAX FORCE AG 5.3 5.9 5.4 5.5 5.4 5.6 (4-AD) 1.1 1.2 1.1 2.0 2.0 1.9 UP SLOW WAIT HJ 2.6 2.6 2.6 1.1 1.2 1.2 MOVE JK 1.6 1.5 1.6 2.3 2.4 2.5 HIT TOP HK 4.2 4.1 4.1 3.4 3.6 3.7 MAX SUPPLY HL 5.0 5.1 4.9 5.0 3.1 5.1 MAX FORCE HM 5.7 5.8 6.0 5.6 5.8 5.5 TOTAL WET LOC5 (4-AD)+HJ 3.7 3.8 3.7 3.1 3.2 3.1 Note: (i) All figures in cells are seconds (s). (ii) Times are the average of 2 old and 2 new feeders respectively. (iii) "New" and "old" mufflers refer to new and blocked Allied Wotan mufflers respectively. (iv) "PA" muffler refers to Portland Aluminium straight through silencer. (v)" TOTAL WET TIME" is the time the plunger is fully extended. During this time there is the greatest risk of plunger corrosion. It is calculated as follows: dwell time - hit crust time + wait time (vi) "Trace sector" is the time between points in the pressure traces Figures 4-3 and 4-4. CHAPTER 4 PNEUMATICS PAGE 156 It was decided to use orifice plates in the air lines which were also considerably cheaper than the other three alternatives ...about $20,000. Installation on the exhaust side of the DCV was unsuccessful due to back pressure causing the valves to stall and by-pass (or stop working completely). Orifices (12mm) were placed in the airlines to increase the movement time from approximately 0.9s to 1.6s. Initially a restrictor was placed in the supply line at the pot (line B in Figure 4-1). On the exhaust line for the downstroke (line A in Figure 4-1), the restrictor was produced by drilling a 12mm orifice in a closed ball valve. In this way, if unsuccessful, the valve could be opened and there would be no back pressure and the feeders would operate close to the original design of fast speed. It was found that there was an annoying noise from the orifices, so another restrictor was placed in both pipes near the DCV. This solved the noise problem. As the cause of anode effects is often difficult to determine, operators go to any length to stop them. It was found that over a period of time some of the ball valves were opened in a (vain) attempt to get better pressure at the crust. In October 1993, a comparison was made of the number of anode effects for pots with valves open to those with valves closed. This showed no difference in anode effects. In November, handles of the ball valves were removed to avoid the temptation of operators opening the valves and causing the feeders to fatigue. The problem with this cheap design is that it meters air into the cylinder as well as out. It is best to only meter out so the piston speed was smooth and has repeatable stroke times. Pressure sensing showed full pressure is attained, but there was an extension to the exhaust time. CHAPTER 4 PNEUMATICS PAGE 157 4.5 RESULTS Some 144 combinations were tested. It would be too cumbersome to explain every one, so only the critical ones are discussed here; a more detailed analysis is in Kissane (76). 4.5.1 Location The feeder at location 1 generally moved faster than location 5 (Table 4-1 and 4-II). Time to hit the crust was about the same and so was the wet time. This cross checks with the observation in the plunger wear investigation that there is no trend of higher wear at any location in the pot (refer Section 5.3.7). Some 60 pots in both lines were checked to see which plungers move the quickest from a rest location. Location 1 was the fastest 85% of the time and location 5 was the slowest 70% of the time. However, there was no trend as to which plunger hit the crust first. The fastest was #1 for 19% and #5 for 13%. Thus, the pattern for movement was not the same as hit the crust. As they hit the crust about the same time, then wet time would be constant and hence plunger wear (from dissolution in bath or friction from hitting the crust). On average, the slowest feeder was 0.5s behind the fastest. Feeder-to-feeder variation in speed was more important than location in the pot. However, as location 5 starts to move the slowest, it will be affected more if the cylinder piston and rod seals have tight clearances or there is drag from assembly component friction. For plants that have bushes to keep the dosing unit tracking correctly, be aware that this may have an effect on stroke speed. 4.5.2 Feeder Age At location 1, older feeders were faster than new feeders by 0.2-0.3s at the original speed and 0.7s at slow speed (Table 4-1 and 4-II, Figure 4-3 and 4-4). At location 5 there was still a 0.7s spread in stroke time for slow speed, but no difference in speed for the original pneumatic design. Hence, the action of slowing the feeders down caused a greater variability in speeds of feeders across the pot. Both the old and new feeders CHAPTER 4 PNEUMATICS PAGE 158 attained full pressure at the same time; only the travel speed into the bath was different at slow speed. 4.5.3 Mufflers Feeders using old mufflers took 0.4-0.8s longer to get to maximum force on the downstroke for the original speed (Table 4-1 and 4-II). They have only a 0.2s effect at the slower speed. Therefore, slower speeds are not affected as much by partially blocked mufflers in time to reach maximum pressure. On the upstroke, the time to reach the top was very similar for old and new mufflers, but location 5 was slower by 0.2s for the original speed. Generally, the old mufflers affected the starting speed on the up stroke as there was a greater waiting time before they received sufficient pressure to accelerate, however, they still reached the top in comparable time. It has been found at Portland that a blocked muffler will cause so much back pressure that the piston will not rise at all. This was verified in 1994 by Tl 1 when it was found that blocking of mufflers can increase the back pressure from lOOkPa to 350-400kPa. Thus, pressure at the crust could halve if a muffler is blocked by a reasonable amount. This may not be obvious to an operator assessing the cause of repeated blockages as the plunger goes up and down "normally". Installation of pressure gauges at each muffler is indeed a solution, but the better solution is to install silencers similar to that designed for Portland. Mufflers may block because of cylinder lubricant exhausting when exposed to excessive temperature. Selection of high temperature lubricant minimises this effect (Section 7.4). For both integrated and independent feeders, the effect of partly blocked mufflers on feed cylinder exhausts is to change shot size, however, the effect is less on the small cylinders used for independent feeders (due to the smaller volume of air vented) and feeders with slow speeds. Note that blocked mufflers also contribute greatly to air CHAPTER 4 PNEUMATICS PAGE 159 consumption. Any blockage on a DCV generally results in the internal spool or poppet not seating correctly and letting air bypass internally through the DCV. At Portland, this was found to be the cause of the largest air leaks on pot feed air systems. This can be detected by a continual hissing noise at the DCV or muffler. Plants can minimise muffler effects by any or all of the following: (i) routine muffler changeout or clean, (ii) using a straight through silencer (such as the Portland design) (Figure 3- 2(b)), (iii) re-routing DCV exhaust into fume duct or pot cavity, and/or (iv) using high temperature lubricant in cylinders. 4.5.4 Speed Change The installation of restrictors increased stroke time by about Is on new feeders but only 0.4s for old feeders (Table 4-1 & II, Figure 4-4). This difference may be due to minor seal leaks that make the piston less sensitive to backpressure. The time to reach maximum force increased by 3 s for location 1 and 2s for location 5 which was much larger than the change in the movement time. This was found to be the case in the potline when the feeders were slowed down in November 1991. Dwell time had to be increased by 2s due to increased frequency of unscheduled anode effects from blocked feeder holes after the speed change. Hence, for any crustbreaker speed change, the dwell time must cater for the change in exhaust as well as supply pressures. The wet time changed by 2.2s for old feeders and 2.5s for new feeders; a rise of 147% and 189% respectively. Hence, restrictors have a very large effect on wet time. Plunger wear was an unfortunate side effect of the speed change. The disadvantage of increased plunger wear would be compensated by dual dwell time or material change (Section 5.6). CHAPTER 4 PNEUMATICS PAGE 160 From original conditions, the Portland speed change dropped kinetic energy by 75% for downstroke and 48% for upstroke for new feeders, and 60% downstroke and 80% upstroke for old feeders. This benefit for upstroke is the main aim of Portland's speed change as this is when the highest deceleration rates are experienced. Following the speed change, fatigue failure of piston rods reduced by 60% (Figure 4-6). FIGURE 4 - 6 PISTON ROD FAILURE HISTORY MECHANICAL FAULT HISTORY REPORT . C-CONTROL CHART Ftradfc>ra Installed between 1 January 1939 and 30 September 1 995 fauHType: 212 1 1D LCL=B.O 0 H •p.. ,. T.., , ,.T , r ( ., r.TT T-[—r i i i i I i | i i i i i ii'i'inttTi I'IIII i r i | i i i i i r-i-fr t i r t i i | i i i i i i i j i I i i r1 r •[ • JAMB9 SEPBB MAY30 JANS1 5EPS1 MAYS2 JAN93 5EP93 MAYS4. JANa5 5EPS5 Dote In aha Ned Subgroup Sizes: n=1 Total number at F&flders In stalled : 1+772 Date Created: 2 October f 995 Total number af fauHs: 40429 Numb«r matching aelaction: 1155 ( 35) 4.5.5 Dual Dwell Time One way of reducing the impact of the longer wet time for slow speed is the use of dual dwell time. A dwell time of 2s is long enough for the alumina to be fully discharged (as shown in the shot size testing experiments in Section 2) but not long enough to hit the crust. For a "hit crust" time of 2.0-2.9s, a 2s dwell time will mean that the plunger will not reach (or barely reach) the crust. The subsequent dwell time (say 4s) every 2-3 CHAPTER 4 PNEUMATICS PAGE 161 cycles ensures that maximum pressure is attained to guarantee a hole. The net effect on wet time is shown in Table 4-IJT. One can see that, even with the slow speed, one can reduce the wet time to the same period as before the speed was changed if a long shot every 3rd time is used. Dual dwell time wet time is 50-70% less than that of a conventional single dwell time system for slow piston speeds. For fast speeds, the effect is 20-70% but it is more difficult due to accuracy of timers for small time intervals. Hence, dual dwell time is more effective for slow speed crustbreakers. It is a very cost effective idea, as the only cost is a computer program change and somefine tunin g to determine the best dwell times. Note that dual dwell time is also useful for independent feeders as these usually operate at slow speeds. The dwell time is already less than for integrated feeders (because they do not haveto allo w for alumina discharge) but the principles are the same. There has been no change in anode effects since dual dwell time was introduced in Portland late in 1992. As will be seen in Figure 5-9, there was also a drop in the wear rate of plungers after the change. 4.5.6 Pressure Required to Break a Hole The pressure required to break a hole can be estimated from the traces by comparing the pressure difference between supply and exhaust at typical plant dwell times. For slow speeds, a pressure of 630kPa for new feeders and 560kPa for old feeders was seen for slow speeds (Figure 4-3). This is effectively 880-890kPa at the crust for a 100mm O.D. plunger. This estimate is probably high as the 4s dwell time used in this calculation is a conservative estimate. However, it does illustrate the problems experienced in blocked feeders in plants using 700kPa plant air supply pressure. CHAPTER 4 PNEUMATICS PAGE 162 TABLE 4 - HI DUAL DWELL WET TIMES. FEEDS NEW FEEDER OLD FEEDER PER MUFFLER DESIGN MUFFLER DESIGN CONDITION CRUSTBREAK LOCATION PA NEW OLD PA NEW OLD TIME TO 1 1 2.7 2.8 2.7 1.9 2.0 2.0 HIT CRUST 5 2.9 2.9 2.9 2.0 2.0 2.1 @ SLOW SPEED DOWN WET 1 1 1.4 1.2 1.3 2.1 2.1 2.0 @4.0s DWELL 5 1.1 1.2 1.1 2.0 2.0 1.9 DOWN WET 1 1 0.0 0.0 0.0 0.1 0.0 0.0 @2.0s DWELL 5 0.0 0.0 0.0 0.0 0.0 0.0 AVE WET TIME 2 1 0.7 0.6 0.6 1.1 1.0 1.0 @2/4 DWELL 5 0.6 0.6 0.6 1.0 1.0 1.0 3 1 0.5 0.4 0.4 0.8 0.7 0.7 5 0.4 0.4 0.4 0.7 0.7 0.6 DUAL/SINGLE 2 1 50 50 50 50 50 50 %DROP 5 50 50 50 50 50 50 3 1 70 70 70 70 70 70 5 70 70 70 70 70 70 ORIGINAL SPEED 1 1.3 1.4 1.4 1.5 1.5 1.4 WET TIME 5 1.4 1.2 1.5 1.0 1.3 1.3 DUAL/ORIGINAL 2 1 50 60 50 30 30 30 % DROP 5 60 50 60 - 20 20 3 1 70 70 70 50 50 50 5 70 70 80 30 50 50 Notes: (i) Figures in cells are seconds (s). (ii) Down wet time = (dwell)-(hit crust time). (iv) Average wet time at 3 cycles/tong shot = [(wet time at 2.0s)+{wet time at 2.0s)+(wet time at 4.0s)]/3 (v) Data is average of two feeders. (vi) Assume dwell times of 2.0s (short) and 4.0s (long). (vii)"% drop" is the percentage reduction in wet time with dual dwell time. CHAPTER 4 PNEUMATICS PAGE 163 4.6 Am SUPPLY OPTIONS Comparison of performance for off-site and on-site tests on the same feeder and stroke times shows much lower wait time for the off-site situation (Figure 4-5). This is due to the fact that there is only one feeder per DCV off-site, versus multi-feeders per DCV on-site. Response time dropped by 0.2 - 0.6s and time to maximum pressure dropped by up to 3s. This illustrates the advantage of the better control for one feeder per valve design. Figure 4-7 shows the response time for one and five feeders operating at one time in a pot for 1.6s stroke time. This shows a faster response time and faster time to reach maximum downstroke pressure for one-feeder/DCV versus 5-feeders/DCV. The results were similar to comparisons of on-site and off-site tests (Figure 4-5). The difference in the one-feeder/DCV off site (Figure 4-5(b)) and on-site (Figure 4-7(b)) illustrates the effect of proximity of the valve to the feeder. The off-site testrig had the DCV only about 2m from the feeder versus 10m on-site. The closer the DCV is to the feeder the quicker the response and the less variable are the stroke time and maximum pressure time. The dwell time chosen for a pot line is generally that of the slowest feeder as dwell time is mostly chosen by people reacting to blocked feeders across the potline. It is usually extended to satisfy the worst pot rather than select a time based on the majority requirement and addressing the problems with the worst pots. Thus, plants generally penalise the good feeders to satisfy the bad ones. This also affects the life and degree of dagging that will occur for feeders that stroke fast. The net effects are plunger wear, dags, cylinder seal failure, and assembly failure. For a multi-feeder pot, what is required is to have all the plungers go down together with little variability across the pot so that the dwell time can be reduced. However, this is not good for the alumina concentration stability of a pot (Section 1.3.13). CHAPTER 4 PNEUMATICS PAGE 164 > > a LL w Q j-iJ-i-.l • j j i i w ,..J.jJ..i 1 ! i ! (9 w LU. Li . JLLLi O fit! tf Li.j-L.1. w i"i * r.L/i PQ 14-] y! .1.. i. —_.-.A ..-- _ I ... t. - > a a UJ •0 o Aia/!*d oz A|Q/!*d 02 M a a. vS A CHAPTER 4 PNEUMATICS PAGE 165 In the pressure traces, point C often shows a "stalling" effect as the piston starts to move. This was better at slow speed (Figure 4-5). This cross-checks with the shot size testing on operating pots where the variability was better for slow speed when changing from 2.5s to a 4.5s dwell time (Table 2-X) and indicates that the pneumatics of the pot are one of the main contributors to shot size variability. Portland has one of the worst pneumatics for a pot. Compared to Til for example, Portland has 5 feeders (not 3), 125mm bore (not 100mm), 508mm stroke (not 356mm), longer pipe length (20%), and larger pipes (32mm not 25mm). This amounts to about 7 times the air volumeto move the pistons full stroke. The capacity of a DCV or valve is rated in a dimensionless unit called CV. The CV rating is a linear relationship to the maximum air flow rate. The higher the CV, the higher the possible air flow rate. Portland has a 32mm valve (versus a 25mm valve at Til) which has a CV of 16 (versus a CV of 14 at Tl 1). Thus, Portland has only a 17% higher air capacity compared to Tl 1 yet it should have a valve 7 times larger to obtain similar air flow characteristics. Clearly, Portland is relatively starved of air supply. Note that the DCV size at Til is similar to that at most Alcoa/Alcan/Alumax plants with equal or less feeders/DCV. Hence, Portland the worst situation of any of these plants. The off-site test rig had a CV of 5. Hence, for the same control on a 5 feeder pot, a CV ha* of 25 would be needed. The Portland valve only 60% of this capacity. This is why the feeder "stalls" on the pot and does not on the testrig of f site. The ideal situation would be one feeder/DCV so that many of the pneumatics problems are minimised. Feeders that utilise separate crustbreaker/feeder have the feed cylinder pneumatically operated directly from a cylinder with no spring and little cushioning as the stroke is only 50mm and the cylinder is only 50-80mm in diameter (Figure 1-12). The cushioning is not a major issue due to the small size and masses. The dosing cylinder valve is also usually positioned directly above the cylinder or close to the cylinder, CHAPTER 4 PNEUMATICS PAGE 166 providing better repeatability as there is no variability due to compression of air in pipes between the valve and the feeder cylinder. Pechiney API8 pots have two hoppers with 2 feeders and 2 crustbreaking units on each (20). The valves for each hopper are on the side of the hopper. The pipe/hose length is about 3 metres for the closest and 4 metres for the farthest crustbreaking unit. (This is about the same as that used on the Portland off-site unit viz. about 3 metres). In the AP30, the valve is on the feeder itself. In the case of Portland operating pots, the farthest cylinder on the five cylinder manifold being controlled by the same DCV is 19 metres from the valve and it takes about one second to pressurise sufficiently to move the valve. The closest feeder is 10 metres from the DCV. This large difference in distance from DCV to feeder reports as a slower initial spool speed and different feed rate. In addition, each plunger/spool may start moving at slightly different times each cycle. The performance of one feeder can affect the variability of the others in the same pot which indicates that plants with 1 or 2 feeders per DCV may be less affected by pneumatics comparedto Portland . Clearly, both pneumatic testing (Chapter 4) and shot size testing (Chapter 2) indicate that the pneumatics of feeder design have a significant effect on shot size accuracy, and that independent feeding is superior to the integrated multi-feeder/DCV design. CHAPTER 4 PNEUMATICS PAGE 167 4.7 GRADUAL INCREASE IN DWELL TIME Note also that the dwell time for a plant needs to be revised objectively from time to time to ensure that the dwell time does not increase gradually over time. Often, if anode effects occur, management will "err on the safe side" by increasing dwell time. This can happen on several occasions spaced over years and people do not realise the gradual change. An anecdote of this gradual trend of increasing dwell time is an experience the author had when visiting a plant with a 4s dwell time when a 2s dwell time would be sufficient to do the job (T9). In the middle of a presentation on optimising feeders, a discussion was held on minimising dwell time. A shift supervisor stood up and apologised for the interruption by saying "I'll be back in a moment after I have dropped the dwell time on my Potline". She halved the dwell time and there was no effect on anode effects. Within a few weeks all potlines on this plant had halved dwell time. Some feeders and mufflers had to be changed as they were in poor condition but were being protected from detection by the long dwell time. The observation was made that everyone noted that the time appeared long, but only when someone outside the normal operating group asked basic questions did people decideto d o something about it. We all need to be aware that when we get close to the process we often cannot see the wood for the trees and getting an outsider to audit the operation can be worthwhile. CHAPTER 4 PNEUMATICS PAGE 168 4.8 MAJOR FINDINGS FROM PNEUMATICS INVESTIGATION (i) If not all feeders are operating in a multi-feeder/DCV pot, the speed of the remaining feeders is faster (which affects shot size for integrated feeders). (ii) The time that the plunger is wet and the speed of the feeder are affected by feeder age and muffler quality. (iii) About half the wet time of plungers is due to the waiting time to begin the upstroke. (iv) Operation of one feeder affects the other feeders for a multi-feeder pot. (v) Partially blocked mufflers cause DCVs to bypass internally and can cause large air leaks. (vi) Piston rod failures reduced by 75% after reducing feeder speed at Portland. (vii) Dual dwell operation saves "wet time" of plungers by 20-70%. 4.9 RECOMMENDATIONS FROM PNEUMATICS INVESTIGATION (i) To minimise back pressure on feeders, consider: - routine muffler changeout or cleaning - straight through silencers rather than mufflers, - redirect DCV exhaust to the fume duct or pot cavity - high temperature lubrication on cylinders. (ii) Plants should examine their air supply system to minimise the variation of the speed and wet time of individual feeders and to optimise air usage. (iii) All plants should consider slow crustbreaker speeds and fast response time for upstroke to reduce seal wear and plunger wet time. In addition, integrated feeder plants should consider dual dwell time and regular high pressure strokes. CHAPTER 5 PLUNGER WEAR PAGE 169 CHAPTER 5 PLUNGER WEAR 5.1 SUMMARY Plunger life varies from 5 months to 6 years over a range of smelters. The causes of wear of plungers was examined by routine measurement of O.D.'s of over 10,000 plungers into and out of the workshop, plus about 2,500 in pot measurements of plungers of various materials, mass, length, and shape, since 1992. Tests on 19 materials were conducted. A large range of factors was detected as being significant to plunger wear, with the main ones being (i) penetration distance into the bath, (ii) cumulative wet time and (iii) material of construction. The wear mechanism of plungers is a combination of erosion and corrosion, with corrosion being the predominant factor. Use of high nickel/chrome materials and less penetration in the bath can reduce wear rates by up to 15 times that of cast iron. Plunger life over 5,000 days (15 years) is possible. CHAPTER 5 PLUNGER WEAR PAGE 170 5.2 INTRODUCTION Plunger wear is one of the most frequent causes of crustbreaker overhaul for both independent and integrated feeders (Table 1-V). If the plunger diameter at the base is .P 4VVS D.UOQB4- 'Yb too small orAtoo short, bridging may occur at the crust and the alumina does not flow into the bath. This chapter will examine the pattern of plunger wear and will provide evidence to support or refute the common theories of the plunger wear mechanism. The results of an intensive experimental program on different plunger materials and designs will be documented. 5.2.1 Pattern of Plunger Wear Plungers are cylindrical in shape with a bottom diameter of 80-100mm. The plunger wears faster at the bottom and generally the shape of a worn plunger is conical as illustrated for a range of feeders removed from the pots at Portland in Figure 5-1. FIGURE 5 -1 TYPICAL WORN PLUNGERS CHAPTER 5 PLUNGER WEAR PAGE 171 In some plants, the shape can have a sudden drop in diameter 10-30mm from the bottom e.g. T9, T15, T16. This is generally due to penetration in the metal pad that causes extreme wear. Plungers made of cast iron can have variable roughness of the original surface but stainless steel bar is very smooth from new. Irrespective of the initial surface condition, the worn surface of a plunger is very smooth. All Alcoa and Alumax plants and most Alcan plants use Alcoa Heat Resistant specification (HR spec) cast iron plungers with beveled end similar to plunger C in Figure 5-2. The bevel acts like a chisel to break the crust. Most other plants use stainless steel bar with flat bottoms. Often plant trials of plunger wear are not concluded due to the long period required to assess results and to poor tracking systems for tracing plungers once they have entered the plant. Little is published on the causes of plunger wear despite a great deal of experimentation by many smelters. The only paper of significance is one by Kaiser in 1984 (54). Estimation of plunger life is very subjective for most plants. Little information is usually kept on plunger life. Life between 5 months and 6 years is quoted, with little data being available to substantiate the statements. One would expect that all feeders and all pots are the same, however, this is far from the truth. Even within a plant there is a wide range of life for individual feeders. 5.2.2 Wear Rate Measurement It is important to specify a parameter that satisfactorily measures plunger wear. Consideration was given to use of mass loss but it was felt that this is impossible to measure on operating pots, and extremely difficult in the workshop as the plunger is generally destroyed in some way in order to separate it from the plunger shaft. Metal loss may give a false indication of bottom edge wear, but is difficult for people to relate to the ability of the plunger to break a hole. A physical measurement is more desirable CHAPTER 5 PLUNGER WEAR PAGE 172 FIGURE 5 - 2 TRIAL PLUNGERS A FLAT BOTTOMMED PA SPEC CAST IRON B HIGH MASS HR SPEC CAST IRON C BEVELLED ALLOY COATED CAST IRON SQUARE SILICON CARBIDE (FIRST TRIAL) SQUARE SILICON CARBIDE (SECOND TRIAL) INCONEL 601 (PIPE) 1. NEW 2. END FALLEN OFF 3. PEELED UPWARDS AND SHAFT EXPOSED TO CORROSION G 80mm O.D. CAST NIRESIST D4 H CAST 31 OSS One of the most important factors in delivering alumina to the liquid bath is a suitably sized hole in the lowest edge of the crust. The plunger diameter O.D. at the bottom edge is the most important feature of a plunger. It was also considered that, whether the attack was corrosion or erosion, the wear will probably be a one dimensional attack, so diameter would be a simple and reasonable parameter. Throughout this report, wear CHAPTER 5 PLUNGER WEAR PAGE 173 rate has been expressed as the "reduction in O.D. of the bottom edge of the plunger over time". This has been measured in mm/month. From 1990, Portland has measured plunger O.D. into and out of the workshop to investigate the wear rate of plungers, and for tradespersons to have a simple gauge to decide if the plunger needs replacing. Over 10,000 measurements have been made from 1991 until the end of 1995. In addition to these routine data, three monthly measurements of plungers in operating pots has been conducted on about 400 plungers every three months from September 1992 to September 1994. 5.2.3 Theory of Plunger Wear Mechanism Possible causes of plunger wear include (a) gas attack when the plunger is raised, (ii) erosion as the plunger passes through the crust, and (iii) chemical attack (corrosion) while in the bath. As there is little evidence of attack of superstructures or the fume system, gas attack was rejected as being insignificant. This leaves erosion and corrosion as the potential causes. Three independent assessments of the cause of cast iron plunger wear were conducted by metallurgical laboratories. Metalab (80) concluded erosion was the cause despite noting no wear grooves on the surface. It was assessed that the Metalab conclusions were based on only a few plungers of unknown life in the pot, so their conclusions were discarded, but their observations were accepted as being valid though limited. CETEC (81) concluded it was corrosion after making similar observations to the previous study. Due to the opposite conclusions from the same types of studies, it was decided to conduct a third assessment, but this time using more samples and doing a CHAPTER 5 PLUNGER WEAR PAGE 174 "blind" assessment where the laboratory was not aware of the history of the samples. Hans Kempe (Kempe Engineering), a metallurgist of some note, sourced Greg Hillard from ETRS as an expert in this type of analysis. Kempe prepared samples for ETRS to do microscopic analyses and chemical analyses of 14 plungers of different batches and ages. Hardness tests were conducted on different locations on the sides and top of the test plungers as it was thought that high hardness should resulti n longer plunger life. The results are presented in Kempe's report (82) and are discussed below. Assessment of hardness can be measured using Rockwell or Brinell measurement units. The Rockwell method uses a spike that presses into the test material and the depth of penetration indicates a rating of hardness. The Brinell method uses a small ball that is forced onto the test material and the diameter of the indentation indicates the hardness rating. It was found that the Brinell method was more repeatable for plunger measurement because the spike of the Rockwell method is affected by the grain boundaries of the cast iron. The Brinell ball tends to be more robust for the location where the test is conducted. Hence, all hardness measurements in this study use Brinell units (BHN). The main observations from these studies were as follows: (i) no microscopic differences between the edge and the middle of the plunger (80,81), (ii) no sign of grooves from erosion (81), (iii) no grain spheroidisation and dissolution of pearlite which would indicate a heat affected zone (82), (iv) comparison of 14 plungers showed wear rates of 1.7 to 6.5mm/month, (v) the longer lasting plungers had molybdenum contents of 4% and copper content of 0.2% (versus 2% and 0.5% respectively for average plungers tested), (vi) hardness was over 200BHN for the better wearing plungers versus 160BHN for the worst (82), CHAPTER 5 PLUNGER WEAR PAGE 175 (vii) high wear rate plungers had less chrome (82), and (viii) there was no difference in hardness between worn surfaces and surfaces that had not contacted the bath (82). The final conclusion was that the cause of wear was mainly chemical attack with flaking off of the reaction deposits occurring on each movement into the crust. Abrasive attack was present but was not considered to be the main factor. This rusting theory was supported at T9 where analysis of the material that fell off some plungers had Fe203 (rust products) on the inner face which previously had been touching the cast iron plunger. To test the corrosion/erosion theories in a simple way, plungers made of Inconel 601 pipe were placed in service at Portland in February, 1991 (Figure 5-2 plungers Fl, F2 and F3). Inconel has a hardness of 130BHN which is about half the hardness of HR cast iron. If hardness is the main factor restricting plunger wear, then the Inconel would not wear as well as cast iron. However, the Inconel proved to have a wear rate of about 14% of cast iron. This clearly showed that corrosion is a dominant factor. Following the metallurgical investigation, a specification for cast iron ("PA spec" for Portland Aluminium specification) was developed by Kempe, Hillard and Kissane (82). This was introduced in January 1992 (Figure 5-2 plunger A). By November 1992, it was not considered that any significant improvement in life was attained by changeover from HR to PA spec, so investigations commenced into alternative materials and operating parameters tofind wha t are the causes of plunger wear and what materials are better than HR cast iron. Figure 5-2 illustrates several of the plungers tested. In hindsight (3 years later), the mistake made in the metallurgical studies was that, although the life of the test plungers was known, there was no knowledge of the pot where the plungers were used. Later it was found that the variability of plunger wear and hardness was within normal variability, so to compare "good plungers" with "bad plungers" was inappropriate as the cause of this large variability was the duty in the pot, CHAPTER 5 PLUNGER WEAR PAGE 176 rather than the material of construction (Section 5.3.2). This is one of the problems of analysis of what statisticians call "happenstance data" where one does an analysis of data that just "happens". It is far preferable to design a controlled experiment where the experimenter knows all the parameters and measures their changes in a log as the experiment unfolds. The reason this was not done at the time was the long time to get data out of plunger experiments and the lack of understanding of what parameters are important in plunger wear. It turned outto be a costly mistake as the PA spec proved to be 40% worse than HR spec (Section 5.6.3). 5.2.4 Procedure for Analysis of Plunger Wear Mechanism If corrosion or erosion are joint causes of plunger wear, then some factors which may contribute to plunger wear would include those itemized in Table 5-1. The effect of these variables can change from plunger-to-plunger both within and between pots. TABLE 5-1 PARAMETERS POTENTIALLY AFFECTING PLUNGER WEAR CORROSION Electrolysis Wet time - Wet time per crustbreak - Crustbreaker frequency - Cumulative wet time Penetration depth in bath Temperature - Bath - Plunger Anode effect control Plunger mass Plunger buildup Plunger material EROSION Plunger speed Crust hardness Plunger hardness Plunger shape Crustbreak frequency CHAPTER 5 PLUNGER WEAR PAGE 177 To determine evidence of plunger wear, extensive trials were conducted on different plunger materials in a range of pot conditions from February 1991. Routine plunger measurements into and out of the workshop (from May, 1991) plus three monthly measurements on operating feeders (October, 1992 to September, 1994) gave early indication of successes and failures on different plunger materials and plunger designs. To establish the factors affecting plunger wear and the relative wear rates of alternative materials tested at Portland, plungers were measured by a pair of calipers each time a feeder came into the workshop irrespective of the cause of failure. This gave a range of life from days to years so a good assessment of wear rates resulted. The tools made to measure plunger diameter on operating pots was a large pair of calipers on poles. A gauge capability study on these tools showed the in-plant tool had a standard deviation of 0.7mm, and the workshop tool had a standard deviation of 0.5mm. This Chapterfirstly examine s why the wear rates of plungers vary so much, then establishes a theory on the wear mechanism. Data from test work have been compiled against plunger wear parameters raised in Table 5-1 and this is discussed in the following Sections 5.3 to 5.6. Following this assessment, in Section 5.7, conclusions are made on the main parameters causing plunger wear. This then leads to opportunities for life improvement and different materials of construction. (This Chapter only discusses plunger life. Section 9.3 discusses the economics of different plunger materials and plunger designs.) CHAPTER 5 PLUNGER WEAR PAGE 178 5.3 VARIABILITY OF PLUNGER WEAR RATE Before one can consider methods to improve plunger life, one must understand the causes of variability of plunger wear. Using mainly Portland data, measurement of plunger wear is discussed below. 5.3.1 Overall Observations Figure 5-3 shows the "shot gun" nature of plunger O.D. at repair for 1100 measurements. Plungers that last a very long time are probably not touching the bath and/or crust, so have no corrosion or erosion. Clearly there is a great deal of scatter of wear rates which is not uncommon with most plants. The causes will be discussed below. FIGURE 5 - 3 PLUNGER DIAMETER VERSUS LIFE REPORT PLUNGER DIAMETER VS LIFE: SCATTERGRAW Date Repaired fram 1 January 1991 to 30 Sept amber 1995 Pari No: 351 A M E T E R 50 CHAPTER 5 PLUNGER WEAR PAGE 179 To gain any understanding of plunger wear, one has to identify the causes of this variability. This has to be done before contemplating changes to material of construction. It is best to optimize what one has, before changing it. 5.3.2 Effect of Single Batch Variability on Plunger Wear Rate It is possible that the cause of variability is at manufacture. It was decided to do a trial of similar plungers and follow their performance in the pots over a period of years. These plungers were examined prior to installation in the pots to establish if they were different, and then assessed the wear rates when they were removed. A set of 60 HR spec cast iron plungers were tested. Cast iron plungers are made in an inductive furnace. Each batch (heat) was poured into a ladle which in turn was poured into a group of sand moulds which contain five plunger shapes. The plungers were coded individually with heat (4 were made), mould (3 were made from each heat) and position in the mould (5 positions). Each furnace heat was analyzed separately using a spectrometer and a potential gauge capability study was held on the spectrometer. This technique is a statistical analysis to determine the accuracy and repeatability of a gauge or measuring process. The spectrometer analyses were statistically "in control" for all elements except silica (which is not considered an important element for wear). Hence, the plungers were identical chemically before service in the pot. It was considered that if erosion was important in wear, then hardness should be an important variable. Firstly, the measuring unit had to be tested for capability and a testing location had to be selected that gave repeatable results. Brinell hardness was tested at many locations on several plungers. It was established that they could be tested at any location on the plunger and would have a reading that was within 20 units (on a 95% confidence level for an average reading of 170BHN). Hardness results for the 60 plungers are represented in Table 5-II. CHAPTER 5 PLUNGER WEAR PAGE 180 TABLE 5-II PLUNGER HARNESS TRIAL DATA HR SPEC CAST IRON MOULD HUSI1IUN 1 2 a 4 b 7 RIAL #1 - HK SPLC BHN WtAPT BHN WtAK BHN WbAK BHN WbAK BHN WEAK 15-May-1991 RATE RATE RATE RATE RATE HbAI MOLL) (mm/m) (mm/m) (mm/m) (mm/m) (mm/m) A 1 217 1.5 207 5.3 229 207 5.8 217 2 197 200 4.4 217 207 2.7 207 2.0 3 207 207 179 1.9 207 0.2 201 B 1 212 2.5 197 3.4 207 4.1 217 197 2 212 217 2.6 207 2.0 223 212 3 212 2.0 197 217 3.8 217 7.8 201 3.4 c 1 197 1.8 217 3.1 223 3.0 207 207 2 197 5.3 197 1.1 197 4.4 191 2.2 201 3 187 2.8 217 207 207 2.6 187 2.7 D 1 217 2.5 212 6.4 217 207 1.6 212 4.1 2 201 2.5 207 212 4.3 217 1.5 212 3.8 3 212 207 4.1 212 3.3 212 217 2.1 Average 206 2.6 207 3.8 210 3.4 210 3.1 206 3.0 Standard Deviation 10 1.2 8 1.6 13 1.0 8 2.5 9 0.9 Minimum 187 1.5 197 1.1 197 1.9 191 1.5 187 2.0 Maximum 217 5.3 217 5.5 229 4.4 223 7.8 217 4.1 Number of data points 12 8 12 7 12 8 12 6 12 6 MOULD No. #1 #2 #3 MOULD No. #1 #2 #3 WEAR RATE (mm/month) HARDNESS (BHN) Average 3.0 3.1 3.3 Average 211 207 206 Standard Deviation 1.1 1.4 1.6 Standard Deviation 8 8 11 Minimum 1.5 1.1 1.9 Minimum 197 191 179 Maximum 5.3 5.5 7.8 Maximum 229 223 217 Number 11 13 11 Number 20 20 20 HR SPEC OVERALL HARDNESS WEAR RATE (BHN) (mm/month) Average 208 3.2 Standard Deviation 9.5 1.4 Minimum 187 1.1 Maximum 229 7.8 Number of data points 60 35 Note: Gaps in the table are due to missing wear rates or wear rates over 10mm/month (which are probably due to errors in dates in pots). CHAPTER 5 PLUNGER WEAR PAGE 181 Tests found that the position in the mould (1,2,3,4,5 on the top of Table 5-Et) and each of the four moulds (A,B,C,D) had statistically similar wear rates. Hardness varied between 180BHN and 230BHN. Hence, manufacture of cast iron plungers was repeatable for hardness. Note that this variability was similar to plungers of different ages tested in the Kempe/ETRS investigation in Section 5.2.3. Thus, the 60 plungers were identical chemically and had the same hardness. Any variability in service would be due to pot variability and not to how the plungers were made. These plungers were placed in service within a two week period and the wear rate was determined when the feeders were removed over the next 3 years. Table 5-II shows that there was no statistical difference in wear rate between heat, mould or position in the mould. The plunger wear rates were variable but not statistically different. Plunger wear rate varied from l.lmm/month to 6.4mm/month (which was similar to the variation seen in samples tested by Kempe/ETRS (82) mentioned in Section 5.2.3). In 1992, a similar trial was conducted using PA spec cast iron (Table 5-ffl). This showed similar results in wear rates and no pattern between heat, mould or position. Once again, the manufacture process was not the cause of plunger wear variability for a cast iron of different analysis. In both trials, the standard deviation was 30-50% of the average for these "identical" plungers. The variability was thus very large. Some plungers could last 5 times longer than others despite being identical when placed in the pot. This suggests that variability from pot-to-pot (such as bath level, temperature, penetration distance into the bath and time in the bath) is the main reason for this plunger wear variability. CHAPTER 5 PLUNGER WEAR PAGE 182 TABLE 5-HI PLUNGER HARDNESS TRIAL DATA PA SPEC CAST IRON MOULD POSITION 1 2 3 4 5 TRIAL #2 - PA SPEC BHN WEAR BHN WEAR BHN WEAR BHN WEAR BHN WEAR 25-May-1992 RATE RATE RATE RATE RATE HEAT MOULD (mm/m) (mm/m) (mm/m) (mm/m) (mm/m) A 1 321 341 5.3 363 4.1 341 331 1.2 2 341 3.9 341 4.9 341 6.0 321 5.3 363 3.9 3 341 2.8 321 321 3.1 331 5.2 341 5.4 B 1 363 331 341 5.8 363 363 2 363 7.9 341 4.1 341 4.9 341 1.5 341 3 375 5.0 341 6.0 341 6.3 341 3.8 341 2.1 C 1 362 341 4.3 341 4.8 341 5.0 341 2 363 331 341 352 341 5.2 3 341 4.1 341 352 7.6 341 5.3 363 D 1 363 3.6 363 3.1 341 3.1 341 4.5 375 6.9 2 341 3.0 321 5.9 363 341 363 3.3 3 363 5.4 321 388 6.8 341 341 8.9 Average 353 4.5 336 4.8 348 5.3 341 4.4 350 4.6 Standard Deviation 16 1.7 12 1.0 17 1.5 10 1.4 14 2.5 Minimum 321 2.8 321 3.1 321 3.1 321 1.5 331 1.2 Maximum 375 7.9 363 6.0 388 7.6 363 6.3 375 8.9 Number of data points 12 8 12 7 12 10 12 9 12 6 MOULD No. #1 #2 #3 MOULD No. #1 #2 #3 WbAK RATb (mm/month) HARDNESS (BHN) Average 4.3 4.6 5.5 Average 348 345 344 Standard Deviation 1.4 1.5 1.6 Standard Deviation 14 13 17 Minimum 1.2 1.5 2.8 Minimum 321 321 321 Maximum 6.9 7.9 8.9 Maximum 375 363 388 Number 12 13 15 Number 20 20 20 PA SPEC OVERALL HARDNESS WEAR RATE (BHN) (mm/month) Average 346 4.8 Standard Deviation 15 1.6 Minimum 321 1.2 Maximum 388 8.9 Number of data points 60 40 Note: Gaps in the table are due to missing wear rates or wear rates over 10mm/month (which are probably due to errors in dates in pots). CHAPTER 5 PLUNGER WEAR PAGE 183 5.3.3 Effect of Batch-to-Batch Variability on Plunger Wear Rate To establish if there was a batch-to-batch variability in wear rate from the plunger manufacturer, the wear rate of a random batch of plungers each month for 20 months was assessed for wear rate. A control chart of these data is presented in Figure 5-4. The batch number relates date, month, and year e.g. 13F90 was made on 13th June, 1990 and 11K91 was 11th October, 1991 (where "I" is not used as it looks like "1"). FIGURE 5 - 4 CAST IRON PLUNGER WEAR BY BATCH NUMBER WEAR RATE (mm/month) 10 10 HR SPEC I PA SPEC 2s DWELL 4s DWELL1 " FAST SLOW ' SPEED (SPEED 1 (JCL .^^yt^^-^^A^^ I LCL _ 1 1 2 0 1 2 2 0 0 2 0 2 0 1 1 2 1 1 0 0 1 0 3 4 3 9 5 3 4 1 8 2 1 4 5 4 6 5 1 5 7 4 4 6 F F G K K L A B B B c E G H J J K L M B B C 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 STANDARD DEVIATION HR SPEC DA CDCJ** - 2s DWELL 4s DWELL ' rM 0rC° FAST SLOW ' SPEED I SPEED j (JCL 1 1 1 2 0 1 2 2 0 0 2 0 2 0 1 1 2 1 1 0 0 1 0 3 4 3 9 5 3 4 1 8 2 1 4 5 4 6 5 1 5 7 4 4 6 F F G K K L A B B B C E G H J J K L M B B C 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 9 0 0 0 0 0 0 1 1 1 1 1 1 1 1 1 1 1 1 1 2 2 2 CHAPTER 5 PLUNGER WEAR PAGE 184 As the 20 month data are inside the control limits, despite having a poor variability, the batches are statistically in control i.e. they are predictable 95% of the time. All the batches have statistically the same wear rates. Manufacture of cast iron plungers has no statistical effect on plunger wear rate. As was found in the single batch trials, the standard deviation is about half the average. The average wear rates of batches of plungers were between 3mm/month and 7mm/month, so the spread of wear rates for individuals was much wider than was experienced for individuals in the single batch trial. This illustrates the compounding effects of within-pot variability and across-pot variability. In the single batch trial, all feeders were installed at the same time so had the same potline conditions viz. current, potline changes in bath level, potline changes in resistance. Here the plungers were installed at different times, so there may be time-related changes in pot process parameters in additionto th e variability seen before. This may explain why individual wear rates were much more variable than in the single batch trial. Hence, trying to optimize plunger material seems a bit pointless until the reasons for the influence of pot conditions are understood and all that is possible is done to minimize this variability. These results also highlight how important it is to control experiments of different plunger materials to be within-pots to minimize factors that may influence the significance of results. The bigger the variability, the harder is to determine whether the differences in wear rate are due to natural variation in the process or to the plunger material itself. 5.3.4 Effect of Diameter on Plunger Wear Rate In order to compare wear rates, it is possible that plungers may wear at different rates over time. Is the wear rate the same when the plunger is 100mm O.D. as it is at 60mm? If there is a significant relationship, one cannot compare wear rates shortly after installation to those when the plunger is well worn. CHAPTER 5 PLUNGER WEAR PAGE 185 To detect if this is important, PA spec cast iron wear rates were extracted from three monthly measurements on operating feeders. To avoid pot-to-pot variability and changing pot conditions, the wear rates of different plungers in the same pot were ratioed to each other, then the data were sorted by the original size (Table 5-IV). For example, consider two plungers (A and B) that were measured at the start of a three month period at 100mm and 50mm respectively. If they were measured again at 94mm and 38mm respectively at the end of the three month period, the wear rate would be 2mm/month for A and 4mm/month for B. The wear ratio A/B would be 0.5. In a similar way the data were sorted and grouped into 90-100, 80-89, 70-79, 60-69, and 50-59mm starting sizes. These were called groups 90,80,70,60, and 50 respectively. Each of these were then sorted by their ratio to each of the other sizes. Hence, a matrix was developed of median wear ratios of "original size" on the Y axis, and "wear ratio" on the X axis. The final matrix is shown in Table 5-IV with a composited graph, which ratios all data to a 90-100mm starting size. A total of 395 wear ratios were used to generate this table. There appears to be a reduction of wear rate as the plunger gets smaller. For example, for cast iron plungers, the wear rate at 100mm cast iron plunger is about twice that at 50mm. The effect for more passive materials is less. A similar analysis has been conducted using Niresist and HR spec cast iron plungers with similar results, however, the wear rate does n6t drop as quickly for the Niresist plungers possibly due to its passive nature as a result of high chrome/nickel content. It is reasonable to assume that there is probably a different curve for each material, with the more passive materials having minimal dependence of wear rate on plunger diameter. A trial of twenty 80mm Niresist plungers (Figure 5-2 plunger G) showed that the 80mm wear rate was 18% lower than 100mm plungers. This reinforces the theory that plunger wear rate reduces as the plunger diameter reduces. CHAPTER 5 PLUNGER WEAR PAGE 186 TABLE 5-IV CAST IRON WEAR RATE MATRIX COMPARISON STATISTIC GROUP ORIGINAL PLUNGER O.D. (mm) O.D. (mm) 90 80 70 60 Median Wear Ratio 90 1.0 80 0.9 1.0 70 0,9 0.8 0.9 60 0.6 0.7 0.8 0.9 50 0.3 0.6 0.8 0.9 Pooled Median Wear Ratio 90 1.0 80 0.9 0.9 70 0.8 0.7 0.9 60 0,6 0.6 0.7 0.5 50 0.5 0.5 0.7 0.5 Sample Size 90 40 80 35 32 70 23 37 35 60 11 20 55 26 50 15 7 16 23 Notes: (i) Median wear ratio compares median wear rate of a Y axis value with an X axis value eg. for cell 70 vertical/90 vertical, the median wear rate of the 70-79mm plungers is divided by the median wear rate of the 90-100mm plungers. (ii) "Pooled median wear ratio" uses the mean wear ratio weighted by sample size for each cell, eg. 60/90 cell is calculated as follows: (0.6 x 11) + ( 0.9 x 0.7 x 20) + (0.8 x 0.9 x 55) + (0.9 x 0,6 x 26) =0.6 (11 +20 + 55 + 26) (iii) Median wear ratios of 395 comparisons of PA spec cast iron plungers of different starting diameters in the same pot for the same 3 month period. POOLED MEDIAN WEAR RATE RELATIVE TO 100mm STARTING WEAR RATE 0.8 — 2 ne 2 0.6- S z o 0.4 Ul s 111 r- 0.2-- 5 iu VC 0.0 90 80 70 60 50 STARTING O.D. (mm) CHAPTER 5 PLUNGER WEAR PAGE 187 Some reasons for reduced wear rate as the plunger gets smaller could be: (i) the plunger surface is farther from the crust so is cooler and the corrosion rate is slower, (ii)the hole is smaller so there is less radiant heat, a cooler surface, and lower corrosion rate, and (iii)the plunger is not affected by erosion due to the wedging shape as it hits the crust. Hence, comparisons of plunger wear should only be calculated when the original O.D. of the two different plungers are comparable; say, within 10mm of each other. 5.3.5 Effect of Between-Pot and Between-Plant Variability on Plunger Wear Rate Consider Figure 5-5 where plungers were placed in the pot at the same time. The life to get to 50mm varied from 450 days for pot 1002 to 250 days for pot 2016. Note that the duty of plungers in these pots is theoretically identical - same line, current, dwell time, and (for 8 of the 10 plungers) the same plunger batch (15K90). This illustrates how much wear varies from pot-to-pot. It is best to compare rates within a pot rather than across pots. It is even more difficult to compare across plants. Here there are many factors that may affect wear rates. These include: number of feeders per pot, type of feeder (independent or integrated), DCV size, crustbreak frequency, pot size, penetration depth, muffler type, dwell time, current, shot size, air pressure, plunger mass, plunger manufacturer quality control, original diameter, anode effect kill strategy. Thus it is unwise to read too much into what one plant can achieve compared to another. One needs to do comparisons within one plant (preferably in the same pot) to avoid these variables. CHAPTER 5 PLUNGER WEAR PAGE 188 FIGURE 5 - 5 POT-TO-POT PLUNGER WEAR VARIATION Pot 1002 Plunger size (mm) 120 HR Spec -*- -i- -4- -i- -4- -i- -4- -U 0 50 100 150 200 250 300 350 400 450 500 Days Plunger No/Location -"-15K90 ~r~09K90 -*-09K90 •""15K90 "*"15K90 Date Installed : 16 Nov 1990 Pot 2016 Plunger size (mm) 120 HRSpec 100 * 80 60 40 20 1 1 1 i 1 —i 1 1 1 50 100 150 200 250 300 350 400 450 500 Days Plunger No/Location "•"16K90 -T-16K90 -*T16K90 -""15K90 -*16K90 On Line November 1990 CHAPTER 5 PLUNGER WEAR PAGE 189 5.3.6 Effect of Within-Pot Variability on Plunger Wear Rate Figure 5-6 illustrates variation within a pot. Pot 2090 had one plunger reach 50mm in 400 days and the rest reach 50mm in 600 days. Pot 4087 had a scatter of wear rates to give lives between 150 and 300 days. Note also the difference of pot-to-pot as well. In the case of 4087, four of the plungers came from the same batch and the best and worst plunger in the pot were from the same batch. Wear rate can be up to 100% different between plungers in the same pot despite being the same original size from the same batch and installed at the same time. CHAPTER 5 PLUNGER WEAR PAGE 190 FIGURE 5 - 6 WITHIN-POT PLUNGER WEAR VARIABILITY Pot 2090 Plunger size (mm) 120 HR Spec 100 * 80 60 - ---^ 40 - 20 1 i 1 1 1 i 1— 1 1 1 1 1 i 0 50 100 150 200 250 300 350 400 450 500 550 600 650 Days Plunger No/Location ""-/1 -r~/2 -"-/4 -X-14F90/6 Date Installed: 22 Aug 1990 Pot 4087 Plunger size (mm) 120 100 Plunger No -•-22B91 +121391 *K'22B!jVt "•-22B91 "*-22B91 Date Installed : 19 May 1991 CHAPTER 5 PLUNGER WEAR PAGE 191 5.3.7 Effect of Location Within a Pot on Plunger Wear Rate Table 5-V shows plungers of HR and PA spec cast iron in the same pot for the same length of time. The wear rate of each was divided by that of location #5 and the locations were compared in the table. Although the average increases towards location #1, the minimum and maximum highlight that this is not significant as the lowest wear rate is in fact a #1 plunger and the standard deviations for each location are 20-60% of the average. There is no statistical evidence to prove that wear rate varies with location. TABLE 5 - V EFFECT OF LOCATION ON PLUNGER WEAR HR SPEC CAST IRON LOCATION 1 2 3 4 5 MEAN 1.3 1.1 1.1 1.0 1.0 STD DEV 0.6 0.3 0.2 0.6 0.0 MIN 0.6 0.6 0.8 0.1 1.0 MAX 2.5 1.7 1.4 2.2 1.0 COUNT 10 10 8 10 18 PA SPEC CAST IFIO N LOCATION 1 2 3 4 5 MEAN 0.9 0.9 1.1 1.1 1.0 STD DEV 0.4 0.4 0.4 0.5 0.0 MIN 0.3 0.4 0.6 0.3 1.0 MAX 1.5 1.5 2.0 1.9 1.0 COUNT 8 7 9 10 18 Note: Values in table are calculated as follows: relative wear rate = (wear rate/wear rate in location 5) e.g. for HR spec, location 1 has a mean wear rate 30% higher than location 5. CHAPTER 5 PLUNGER WEAR PAGE 192 5.3.8 Recommendations for Data Analysis of Plunger Wear Rates It has been seen that there is a large variation in plunger wear data due to across-pot and across-plant variation. In this type of situation, assessment of different materials cannot be made in a statistical comparison unless (i) the mean wear rates are different by several hundred percent, or (ii) the sample sizes are very large. Clearly, as it is desirable to detect small differences in wear rates of different materials and with the high cost of different exotic materials, one cannot tolerate very large sample sizes in experiments of different plunger materials. Several recommendations are made in order to reduce the sample size. (i) Sample period As seen in Section 5.3.2 and 5.3.3, wear is affected by pot condition. This variability is minimized by comparisons in the same pot over the same time. It is not possible to avoid within-pot variability (Section 5.3.6), but it does not matter which locations one chooses as it has been established that location is not significant (Section 5.3.7). To avoid short term changes to pot conditions or timing errors, a period of 3 month intervals is wise. The 3 month period gives a trend relatively quickly, so that one does not need to wait years to see the effect of different designs. (ii) Original O.D. As mentioned in Section 5.3.4, plunger wear reduces as the plunger gets smaller. Hence, comparative trials must be between plungers of similar original O.D. e.g. within 10mm. (iii) O.D. Measurement Accuracy Comparisons of diameter will be affected by the accuracy of the measurement. From gauge capability studies, it was established that a wear rate should not be used for mean (average) wear calculations unless there was a diameter reduction of at least 5mm. CHAPTER 5 PLUNGER WEAR PAGE 193 However, median wear rates can use all data because medians are not affected by extreme values in a population. For example, averaging a sample of nine values at 0.1 and one at 3.0 gives a mean of 0.4. A better estimate of this sample would be 0.1 (which is the median). Medians are more robust to outliers than means, so will be used in future for comparisons. Most comparisons also table means for information purposes. (iv) Sample Size For statistical significance calculations, it is preferable that the sample size is at least 8; the higher the better. Standard deviations can be used for values of 8 or more. Standard deviations are a useful measurement of plunger wear variation as the populations of the sample sizes generally have a normal distribution. Hence, the following conditions are recommended for plunger wear data of different plunger materials or feeder conditions: (i) compare plungers in the same pot over the same time (3 months minimum), (ii) plungers to be within 10mm O.D. of each other at the start of the trial, (iii) use medians, and (iv) use sample sizes of at least 8 and calculate standard deviations. 5.4 TESTING CORROSION THEORY It is unlikely that electrolysis is a major factor in plunger wear as it would be expected that the whole bottom end of the plunger would dissolve. This would mean a reduction in length. This does not occur on plants except those with insulation breakdown and concurrent long periods of the plunger in the bath. Plants with arcing or with deep penetration into the metal are the only cases detected that could possibly have electrolysis problems. This source of corrosion is therefore not discussed further, except in Section 8.7 where arcing is discussed. CHAPTER 5 PLUNGER WEAR PAGE 194 5.4.1 Wet Time Given that plungers generally cycle regularly in the bath, any time that the plunger is in the bath there is more risk that the temperature and shedding of rust layers will be aggravated by immersion time. Consider the effect of wet time which comprises (i) wet time per crustbreak, (ii) frequency of crustbreak cycles, and (iii) cumulative wet time viz. (i) x (ii). 5.4.1:1 Wet Time per Crustbreak Dwell time is defined as the time that the air extends the crustbreaker to when it retracts the crustbreaker....DCV "on" to DCV "off. It includes the reaction time of the valve, down stroke time and "wet time" (where the plunger is potentially stationary in the liquid). For the purpose of this discussion, wet time will be considered as the time in the bath priorto DCV operation to retract the plunger, and assume that the retract delay time is a fixed ratio to the dwell time. This may not be exact, but it is probably a reasonable assumption. It will be more inaccurate for pots with 3-5 feeders than one-feeder/DCV pots as the delay time is greater due to large air lines. Portland, Til, D12 and T7 may show conservative wet time values as times may be up to double that indicated for these multi-feeder/DCV pots. For most situations, the reaction time is consistent if the valve is in good condition, there are no air leaks and there is no back pressure. Stroke time is consistent if there is constant friction on the drive mechanism and cylinder seals. Hence, the main criterion considered in most plants is changing dwell time to vary the wet time. Section 4.6 shows, for pots with multiple feeders per DCV, other factors can affect wet time. Each feeder and pot may have quite different response characteristics due to DCV efficiency, CHAPTER 5 PLUNGER WEAR PAGE 195 number of feeders on line, air line length and feeder type. Thus, there are potential errors introduced by using dwell time as wet time. Figure 5-7 shows the plant average wear rate for different wet times per crustbreak for similar plunger specification for 13 smelters. There appears to be a general trend of reduced wear for less wet time. FIGURE 5 - 7 CAST IRON PLUNGER WEAR RATE VERSUS WET TIME PER CRUSTBREAK WEAR RATE (mm / month) T9 7 at ~m FEED ON ANODE EFFECT A DO NOT FEED ON ANODE EFFECT 6 T6(1) • T6(3) 5 • 4 3 - T6(2) D12 T11 1 T16 2 T3 • £6 1 T2 Y • f) 1 1_ —I —I 1 1 0.5 1.0 1.5 2.0 2.5 3.0 WET TIME / CRUST BREAK ( s ) Note: Cast iron specification is the same at all plants and T1 is Portland. CHAPTER 5 PLUNGER WEAR PAGE 196 The requirement of selecting a dwell time is to have sufficient time, after the normal stroke time, to cater for such things as: (i) alumina discharge (for integrated feeders), (ii) piston seal leakage, (iii) variability in mains pressure, (iv) crust hardness variability, (v) DCV condition, (vi) cylinder friction from new seals, and (vii) air leaks. For pots that have multiple feeders per DCV, the dwell time is also related to the historical slowest feeder on the manifold (Sections 4.6 and 4.7). Independent feeders do not have to wait for alumina discharge so have lower dwell time and wet time compared to integrated feeders. They only need to reach full pressure. As the DCV is usually close to the feeder itself (within 2m), the response time is much faster than for an integrated feeder which may have pipe lengths of 20m (Section 4.6). integrated feeders usually have a wet time of 1.2-2s versus about 0.2s for independent feeders. Clearly there is potentially a greater effect on dagging and plunger wear for integrated feeders. Tests at T9 have shown that very long dwell time does not necessarily increase the chance of breaking a hole. Experience in each plant will determine the optimum dwell time, but results at T2 and T3 suggest times of about 1.0s are sufficient after completion of the stroke for integrated feeders. From A3 tests at Portland, independent feeders could be retracted within 0.2s of the end of stroke. The reduction of dwell time is usually at zero cost, with potentially significant improvement to plunger life. Following recommendations by the author, all Alcoa smelters reviewed dwell time. T9 and T15 halved dwell time in 1991/92, with many CHAPTER 5 PLUNGER WEAR PAGE 197 other plants reducing by 10-30%. Most plants found that the long dwell times were covering other errors in the hardware...such as blocked mufflers, air leaks, deteriorated air hoses, faulty valves. All plants have improved feeder life since 1992 (Section 10.3.2). Another observation of time in the liquid is the pattern of wear. Plants that have a very long stationary time (T9, T15, T16) may experience a "bottle" shape wear which is often referred to as a "short" plunger. This shortening is one of the major causes of removal (Table 1-V). Plants with less stationary time each crustbreak show a more uniform conical shape. Short plungers may be due to rapid corrosion from the end of the plunger penetrating the molten aluminium in the pot. FIGURE 5 - 8 CAST IRON PLUNGER WEAR RATE VERSUS CRUSTBREAK FREQUENCY WEAR RATE (mm / month) T9 7 • m FEED ON ANODE EFFECT 6 T6J1) A DO NOT FEED ON ANODE EFFECT T6(3) 5 4 T6(2) D12 3 • T4 • • T15 T1 T11 A 2 • T"6 T3 A17 T2 1 0 .. J I 1 I. 1 -J 1 1 1 1 100 110 120 130 140 150 160 170 180 190 200 CRUST BREAK FREQUENCY ( s ) Note: Cast iron specification is the same at all plants and T1 is Portland. 5.4.1:2 Crustbreak Frequency Figure 5-8 illustrates the effect of crustbreak frequency of wear rates of 13 smelters. Despite the variability from plant-to-plant, there is a general trend that the less frequent CHAPTER 5 PLUNGER WEAR PAGE 198 the breaks, the lower the wear rate. This could be relatedto either erosion or corrosion (or both) An example of the effect of crustbreak frequency is for pots that were converted from AEDDto A2 feeders. Figure 5-9 shows the change from a normal AEDD operation to an A2 feeder operation when the frequency of feed is 50% more frequent than the AEDD dueto a smaller shot size. The AEDD would have reached 50mm in about 400 days, yet the A2 would have failed at about 250 days - roughly a similar ratio to the shot size ratio between normal and A2 feeders. FIGURE 5 - 9 EFFECT OF CRUSTBREAK FREQUENCY CHANGE ON PLUNGER WEAR RATE Pot 2009 Plunger size (mm) 120 A2 Inst. 26-M1 HR Spec 1005 AEDD ^"""^^Nllt>^r~" ^^s^v^ A2 6.0mm/month 80 AEDD 3.8mm/month 60 40 - 20 . 1 1 1 1 -I 1 1 u 1 50 100 150 200 250 300 350 400 450 500 Days 22B91/1 01O91/2 "^"l8B91/3 22B91/4 -* 22B91/6 Date Installed 25 May 1991 This increase in wear rate was a similar effect to the 25% increase in wear rate when there was a 25% reduction in shot size on installation of spool inserts. CHAPTER 5 PLUNGER WEAR PAGE 199 Figure 5-10 is a "box and whiskers" graph of plunger wear rate versus month repaired. This type of graph shows a box with "tails" at top and bottom with a horizontal line inside the box. The "whiskers" indicate the highest and lowest value of the population. The top (and bottom) of the "box" are the values that have 25% of the population greater than (and less than) that value. The horizontal line in about the middle of the "box" is the median value in the population. After July 1992, there is an increase in wear associated with the introduction of inserts. Median wear rates increased from 4mm/month (between points 2 and 3) to 5mm/month (between points 3 and 4). Thus, there appears to be a general relationship between the crustbreak frequency and wear rate. Safji»L Q UJ H CO LU H Q LU or W CO h- ion 0_ -J LaU: CO -J CO z £ H ro CO haa* CO fK LO III H LL O LU o aCdD _z1 "/J 1 D n 0- CO c irno PQ LuU _i § o a. z as 5 CO OS UJ o CO I- Q _i LU -I p LU Uj -J -a D- § CM CO g c Ig "Cl a) CO Q 0) ID O CM CL r-- 5 ffl (J J rl fi « ») it M 6 00 Ul N a- T- T- T- t- 5 • O I- CKO-H-B v££\I OC-.Er-% CHAPTER 5 PLUNGER WEAR PAGE 201 5.4.1:3 Cumulative Wet Time When crustbreak frequency and wet time per crustbreak are jointly considered (Figure 5-11), the pattern of plant plunger wear rates versus cumulative wet time shows a clearer relationship than wet time/crustbreak (Figure 5-7) and crustbreak frequency (Figure 5-8). The correlation coefficient (r2) was 0.7 which means that 70% of the variability of wear rate can be explained by cumulative wet time. Considering the graph is based on "whole of plant" data, this is a good correlation given the circumstances. FIGURE 5 -11 EFFECT OF CUMULATIVE WET TIME ON CAST IRON PLUNGER WEAR RATE WEAR RATE (mm/month) 8 - 7 FEEDONA.E. "T9 6 5 FEED ON A.E. 4 3 2 T3 B 1 T2B 0 1 10 15 20 25 30 35 40 TIME (min/day) Note: Cast iron specification is the same at all plants and T1 is Portland. This relationship applies to all plants except T19, T15 and T16 which achieve about three times better wear rate than the others. This is probably due to the fact that these plants do not feed on anode effect when there is elevated temperature and major mixing in the pot which would lead to higher corrosion. CHAPTER 5 PLUNGER WEAR PAGE 202 5.4.2 Penetration Distance in Bath Table 5-VI shows the penetration of plungers into the bath for average metal and bath depths with new cathodes in 22 plants. It is indicative only, as each pot in a plant can be different depending on liquid level, distortion of the cathode, or erosion of the cathode. It gives a general trend of what penetration distances exist in different plants. TABLE 5-VI PLANT PENETRATION COMPARISON BATH METAL PENETRATION PLANT LINE DISTANCE TO CATHODE (mm) (mm) (mm) (mm) T13 All 356 150 100 -100 T11 All 273 158 135 -80/+20 T17 All 273 170 65 -38 D13 All Vary bath level for penetration <0 T3 All 284 190 100 6 T15 8(P100) Vary bath level for penetration 12 T2 All Estimate 15 T6 3 274 190 100 15 D7 Prebake Move feeder vertical <20 T16 All 228 159 90 21 T6 2 268 190 100 21 D3 All 295 185 150 40 T7 All Estimate 30 D1 All Estimate 36 T1 All 295 172 165 42 T10 All Estimate 50 D12 All 260 170 150 60 T15 Vary bath level for penetration 64 T6 1 226 190 too 64 T4 All 250 200 120 70 T8 All Estimate 125 D9 D,E 210 175 190 150 D2 L4 210 200 190 150 D10 All Estimate 150 T9 P100 New Fixed 183 76 140 T9 T51 New Fixed 183 76 183 Note: "Fixed" refers to the plunger is at a fixed height relative to the base of the pot for all pots. These plants target different bath levels for different penetration depths. Penetration varies from -100mm to +183mm showing the wide diversity of operations. If plunger wear does vary with penetration, there are many opportunities for plants to improve plunger life by reviewing this parameter. CHAPTER 5 PLUNGER WEAR PAGE 203 Many plants with minimal penetration recognise its importance and targettotal liqui d level height so that the plunger does not get wet (or only by a minor amount). These plants generally achieve very good plunger wear rates. Hence, it is possible to run a plant on virtually zero penetration (on average). Based on these data, trials were conducted of different lengths of plungers at Portland in 1993. Initially a few plungers had 50mm removed from the bottom edge before installation. Then a number of pots with the lowest liquid base were converted to shorter plungers. No problems were experienced with blocked feeder holes. Comparison of different plunger lengths in the same pot showed that 50mm less penetration reduced the wear rate of PA spec cast iron by 60%, and 70mm less penetration reduced the wear rate by 80%. Similar improvements were achieved for other materials as well. For example, 50mm shorter plungers improved life of cast 310SS by 150% and 304SS by 30% (Section 5.7). These results showed very clearly the importance of penetration on plunger wear. Even zero bath penetration often still breaks a hole as the crust bottom edge is higher than the liquid bath. Therisk i s that bath level control may become unstable, liquid levels will drop and plungers do not reach the crust. Each plant needs to explore zones of safe operation over a long period of time to determine the optimum penetration depth. Portland has seen at times a relationship between high anode effects and low liquid level. One needs to control the bath to avoid blocked feeder holes. It was decided to further reduce penetration at Portland from July 1993 by reducing penetration from 43mm to -7mm by using 50mm shorter plungers. This change has had no effect on blocked feeder holes, thus verifying the tests. Together with the change in length, the plungers were changed to cast 31 OSS. These changes had excellent results with 8 times the plunger life; a rise from 360 days to 3,000 days (Figure 5-12 and Figure 5-13). CHAPTER 5 PLUNGER WEAR PAGE 204 FIGURE 5 -12 PLUNGER WEAR RATIOS (RELATIVE TO PA. SPEC CAST IRON @ STANDARD LENGTH) WITHIN-POT DATA (a) STANDARD LENGTH (30Snim) PA SPEC INCONEL HR SPEC NIRESIST HI CHROME (25-27%) 310SS INCONEL 304SS INCOLOY (b) SHORT LENGTH (255mm) PA SPEC CAST 304SS BAR "• •y^-yy "•<•• • & 310SS CAST JKS CAST NICROFER BAR WORKSHOP DATA (a) STANDARD LENGTH (305mm) PA SPEC HR SPEC HR SPEC (HEAVY) NIRESIST HI CHROME (27%) HI CHROME(25%) 31 OSS INCONEL 304SS INCOLOY (b) SHORT LENGTH (255mm) 304SS 31 OSS JKS NICROFER WEAR RATIO (relative to PA Spec cast iron @ standard length) CHAPTER 5 PLUNGER WEAR PAGE 205 FIGURE 5 -13 CALCULATED PLUNGER LIFE (RELATIVE TO PA. SPEC CAST IRON @ STANDARD LENGTH) WITHEV-POT DATA (a) STANDARD LENGTH (305mm) PA SPEC INCONEL HRSPEC NIRESIST HI CHROME (25-27%) 310SS INCONEL 304SS INCOLOY (b) SHORT LENGTH (255inin) PA SPEC 304SS 31 OSS JKS NICROFER WORKSHOP DATA (a) STANDARD LENGTH (305mm) PA SPEC HR SPEC HRSPEC (HEAVY) NIRESIST HI CHROME (27%) HI CHROME(25%) 31 OSS INCONEL 304SS INCOLOY (b) SHORT LENGTH (255mm) 304SS 31 OSS JKS NICROFER 1000 3000 5000 6000 PLUNGER LIFE AT PORTLAND (days) CHAPTER 5 PLUNGER WEAR PAGE 206 From early 1994, additional tests were initiated on plungers which were a further 50mm shorter than the previously cut plungers. Hence, these plungers had an average penetration of-57mm for a potline. The trial pots had a penetration more like -100mm as these were chosen as being the lowest liquid level pots in the potrooms. There were no problems with these pots for many months, but by the end of 1994, it was considered that it was too risky to expand to full pot operation as it was likely that the liquid level in all pots may need to be dropped in the future and there may be a danger at some time in the future that there may be problems with blocked feeder holes. Hence, the trial was stopped. The trials of reduced penetration proved that crustbreakers with negative penetration would operate successfully provided that the net distance between the plunger and the crust was not going to drop over time e.g. if surface of the top of the cathode (bottom of the pot) eroded. If the cathode generally rose due to the cathode "heaving" (swelling upwards), then negative penetration would be a low risk method to improve plunger wear rate. A further attraction of less penetration is a reduction in plunger temperature due to the greater distance that the plunger tip is from the hole. As will be discussed in Sections 6.3.6, every 100mm of reduced penetration reduces the temperature of the plunger and crustbreak cylinder by about 70°C. Cooler cylinders improve seal life which is one of the main causes of crustbreaker cylinder failure. 5.4.3 Bath Temperature Chemical corrosion is generally accelerated by temperature increase in either the metal or the liquid/gas. A study of the temperature of the plungers at Portland showed a range from 60-330°C existed (Table 5-VTI and Section 6.5.6). There was a median temperature drop of 100°C at the plunger for 100°C drop in the pot, but the variability hole-to-hole (S.D. 500°C) was so large that significance could not be seen across the pots. In addition, the plunger temperatures for a single pot changed randomly over a period of 3 weeks (Table 6-ffl). Most plants have pot temperatures averaging 950- 960°C, but temperatures as low as 920°C and as high as 1,000°C are possible. t- © rs CO h- CO QnN0)0 OT-OOC*>T- III o in J5 CO to $ £! CN^OOOJCO <6 & 3l M S8 E S 111 -c =0> S 0. 1-2 in ;? o o ° in nmcNNN t88 E CD t ^ o> 93 m tOLoaicnoo . M go w ' ui x CL i- gco 2 ce CO 0> ^- ,-, CD _. CO o IO IO CO ui CC O p p (9 O D- cu O 2 D CO ^ (OCM ^ in CN CO 1^ CD CN -J r- CO § 82 Ills Z E < DE V 3 D MEA N MI N MA X NUMBE R ST D LOCn l LOCn 2 LOCn 4 10 LU L0Cn 3 L0Cn 5 w 5 -J CO •* CM O CO •* CM O pa H CD ^f CM O o N o in •* n cj IN om ^n«- 3 0 £ 3 CHAPTER 5 PLUNGER WEAR PAGE 208 The frequency of high temperature periods for pots will cause extra corrosion. Hence plants with large temperature variability may experience worse plunger life. Plunger life at Portland was graphed against bath level but no trend was evident as there was a great deal of variability in the data. 5.4.4 Plunger Preheat Bath temperature affects the surface temperature of a plunger but has little effect on conduction up the shaft (Section 8.4.2:2). The temperature of the plunger affects dags (Section 6.5.7). If the plunger is hot prior to entry into the bath, wear rate may be accelerated. Surveys of temperatures near the plunger hole showed a range of 60°C to 330°C (Table 5-VII). The hotter the pot, the hotter the plunger. There was a large variation in individual plungers. This variation in preheat adds to the variability of plunger wear within and between pots. The temperature will change depending on the percentage of gas that vents under the plunger. The less holes open, the hotter the plunger. The shape of the plunger may also affect plunger temperature. A flat bottomed plunger will heat up more than a conical worn plunger as the former will tend to create more turbulent air at the tip which will promote heat transfer. The effect of this variability depends on the type of plunger material. A highly passive material such as 31 OSS or Nicrofer will probably be affected less than less passive materials like cast iron. As preheat changes by location, pot, pot design and pot condition, no reliable testing has been conducted to identify the significance of preheat. It is concluded that plunger preheat has some effect on plunger wear. CHAPTER 5 PLUNGER WEAR PAGE 209 5.4.5 Anode Effect Control Strategy Anode effects occur when the alumina concentration drops below a bottom limit (Figure 1-6). The bath starts to break down, large high resistance bubbles occur below the anodes and the bath violently agitates. During this time, entry of the plunger into the liquid will result in high corrosion rates. Agitation tends to remove any corrosion products quickly, leading to exposed fresh metal. The higher temperature is an enabler of high corrosion rates. Anode effects can be controlled during this high agitation period by two methods (or a combination of both). (i) feeding the pot (to increase the alumina concentration), or (ii) push the anodes into the bath (to squeeze the bubbles out). The former will result in high plunger wear rates as generally several rapid feeds are necessary. The latter will result in possible bath wash onto the top of the anodes which may dissolve the cast iron connections where the rod meets the carbon anode block. Iron dissolution will drop the quality of the aluminium product which is most undesirable. The effect on metal quality may be small if old anodes are sufficiently large, if the cast iron is properly poured into the join or if carbon paste is used as a joint rather than cast iron. Each plant has a different philosophy on anode "kill" strategy. Figure 5-11 shows that plants that do not feed on A.E. and have similar stationary time have much longer life - about three times that of other plants that do feed on anode effect. The cost of plunger wear is not as significant as the benefit of killing anode effects quickly, however, plunger life will suffer if it is decided to feed on anode effect. There are many plants that achieve very good anode effect frequency and do not feed on anode effect e.g. Pechiney (0.1-0.3 anode effects per day) and D3 (0.2 anode effects per day). Thus it is concluded that plants should consider the possibility of not feeding on anode effect in order to maximize feeder life and balance this against the risk to metal quality. CHAPTER 5 PLUNGER WEAR PAGE210 5.4.6 Plunger Mass In February 1990, some 20 "long" HR cast iron plungers were installed in several Portland pots (Figure 5-2 plunger B). These had the same distance from bottom edge to cathode, however, they were taller. They had the gap between the top of a normal plunger and shaft collar filled up (Figure 1-11). This increased the mass by 20% yet did not change penetration in the bath. Of the 12 plungers traced, average wear rate was 4.7mm/month with an S.D. of 1.5mm/month. HR cast iron plungers placed in service at the same time averaged 5.5mm/month with a S.D. of 2.7mm/month. There was insufficient evidence to suggest they were statistically different. Hence, it was concluded that mass of the plunger does not affect wear rate. 5.4.7 Plunger Buildup It seems reasonable that a coating of passive or hard material on the plunger will act as a protective layer from corrosion (or erosion) as there is no flaking effect of scale (or abrasion) against the crust. It is difficult to assess the effect of this variable as plunger buildup (dags) come and go almost randomly (Chapter 6). It was attempted to simulate the relative effect of erosion and corrosion by immersing the plunger totally in the bath. In this way, there was no cycling through the crust. By comparing the wear rate to that of normal wear in the pot previously, it was expected that one could estimate the effect of corrosion relative to total wear (erosion and corrosion). Two PA spec cast iron plungers were immersed continuously in the bath at a penetration of about 75mm. The plungers were withdrawn, measured, then immediately immersed again. Initially this was done hourly, but as there was little to no wear, the practice was increased to daily. This continued for 8 days; it was quite surprising that the wear was so slow. The trial was conducted on another two occasions on different pots with the same results. CHAPTER 5 PLUNGER WEAR PAGE 211 The wear rate was about 10% of the wear rate of plungers in the same pot previously...0.4 mm/month versus 4.5mm/month. (Wear rate was calculated based on an equivalent wet time of 2s/cycle and 160s/cycle.) This experiment at first appears to give results that are opposite to the trend of Figure 5- 11 which suggests corrosion is 70% of wear rate. The plungers in this experiment may be affected by no flaking from the outer boundary layer that occurs with normal plungers that retract out of the hole each cycle. T9 has analyzed the inner layer of the material that falls off a cast iron plunger to be iron oxide...rust. In this Portland experiment, this rust layer was not removed each cycle so the surface of the plunger was protected from further attack. This type of experiment was not effective in simulating plunger wear, but it does illustrate that excessive times in the bath does not have any short term deleterious effect on dissolution of a plunger. Note that long periods in the bath will cause heat conduction up the shaft to the cylinder seals which may affect seal life (Section 8.4.2). It is important during feeder changeout that the plunger is not dropped into the bath to avoid this conduction effect. The same problem arises for new feeders. Pechiney and D3 have a bolt at the base of the plunger to secure it. Portland uses wires to secure the plunger on new feeders. 5.5 TESTING EROSION THEORY 5.5.1 Plunger Speed Portland changed the speed of its feeders in November 1991 but, at the same time, increased dwell time due to the changes in pot pneumatics. Because of the two changes occurring at the same time, it is difficult to separate the two effects. However, if one makes the assumption that wet time has a linear effect on wear rate, then the frictional effect can be calculated. CHAPTER 5 PLUNGER WEAR PAGE 212 Wet time increased from about 1.4s to 3.3s; a 2.4 factor change. It would be expected that the net effect of the wet time change would be 2.4 times the wear rate (using data from Tables 4-1 and 4-H). This would be offset by a reduction in frictional wear against the crust which would be expectedto have a squared relationship between final velocity and wear rate. The peak velocity became 0.6 times the original velocity when stroke time increased from 0.9s to 1.6s. If friction is a linear relationship to the cube of velocity, the friction change should be 0.2 of the original. The net effect of the 2.4 times change in wear rate from wet time and 0.2 times change due to the speed change would be (2.4x0.2) = 0.5; a drop of 50%. Comparison of plant wear rates before and after the change is presented in Table 5-VJTI. TABLE 5 - VIH EFFECT OF PLUNGER SPEED AND DWELL TIME ON PLUNGER WEAR RATE CAST IRON STROKE OVERALL NEW POTS OLD POTS DWELL TIME ALL DATA SPEC SPEED MEDIAN ONLY ONLY HR Fast 2s 5.5 5.0 6.6 5.3 HR Slow 4s 4.8 3.5 4.0 4.9 Statistically significant (90%)? Yes Yes Yes No Difference (%) 14 30 39 Note: All values are wear rates in mm/month for plungers from production pots. Most of the data indicate that the combination of slow speed/long dwell has improved wear rate. The actual effect on median wear rates was a drop of 70% which suggests that the frictional effects (speed affected) may be r^oce influential corrosion effects A (wet time effect). (The exception to this comparison is for new pots which usually run at higher temperature and higher bath levels than old pots, and the holes are generally much larger. Hence, the main wear mechanism for plungers in new pots is corrosion, as erosion is minor.) CHAPTER 5 PLUNGER WEAR PAGE 213 5.5.2 Crust Hardness Hardness of the crust is generally assessed by the fluorospar content of the bath. The lower the fluorospar (CaF2) level in a pot, the harder the crust. No comparative testing of crust hardness has been conducted in this study and no published articles refer to the effect of fluorospar content on hole breakage or plunger wear. This is not expected to be a big variable at the 5-6% fluorospar levels at which smelters usually operate. 5.5.3 Plunger Hardness The effect of plunger hardness is best illustrated by controlled experiments of different plungers as part of this investigation at Portland. Comparison of different materials (which shows analysis, hardness and relative wear rates) is shown in Table 5-EX. As discussed in Section 5.2.3, single batch trials were conducted at Portland using HR and PA spec cast iron. It was expected that, if erosion is a factor in the wear of plungers, a harder material would be beneficial. The hardness of the PA spec plungers was 346BHN versus 208BHN for HR spec. PA spec has a statistically much higher hardness, yet it made no difference to wear rates viz. 1.2-8.9mm/month versus 1.1- 7.8mm/month respectively. The variability was very large for both materials in these tests as these were scattered over many pots. None was tested in controlled experiments in the same pot for the same time. However, Section 5.6.3:1 does show results in controlled experiments that there was a 40% drop in life for the harder PA spec compared to HR spec which supported these plant trials. Trials of Inconel pipe showed wear rates 7 times lower than cast iron despite having a hardness of 130BHN versus 170BHN for HR spec and 350BHN for PA. However, trials of high chrome cast iron with a hardness of 450BHN and 600BHN showed wear rates 2.4 times less than PA spec. These conflicting results suggest that hardness is not a significant indication of plunger wear rates. CHAPTERS PLUNGER WEAR PAGE214 TABLE 5 - IX PLUNGER MATERIAL COMPARISON BRINELL COST/ COST/ WEAR RATE RELATIVE EXPECTED ALLOY 1\SSUM E CAST LENGTH O.D Ni Cr HARDNESS PLUNGER CONVERT RELATIVE TO LIFE FOR PLUNGER LIFE /BAR (mm) (mm) (%) (%) (BHN) ($) ($) 05mm PA SPE WALL O.D. months) (days) CAST IRON HR SPEC (COMPLETE) CAST 305 100 0 1 170 45 1.4 1.4 17 504 PA SPEC (COMPLETE) CAST 305 100 0 1 350 40 1.0 1.0 12 360 PA SPEC (COMPLETE) CAST 253 100 0 1 350 40 16 16 19 576 HIGH CHROME 27% Cr (COMPLETE) CAST 305 100 0 27 600 100 2.4 2.4 29 864 25% Cr (COMPLETE) CAST 305 100 2 25 450 100 2.4 2.4 29 864 304 COMPLETE BAR 305 100 9 19 155 255 3.3 33 40 1188 BIMETAL TIP BAR 305 100 9 19 155 110 52 3.3 3.3 40 1188 BIMETAL TIP BAR 2SS 100 9 19 155 80 52 42 4.2 50 1512 310 COMPLETE CAST 305 100 20 25 120 2.8 2.8 34 1008 COMPLETE CAST 255 100 20 25 80 8.3 8.3 100 COMPLETE ASSUME BAR 305 100 20 25 170 315 5.6 5.6 67 NIRESIST D4 COMPLETE ASSUME CAST 305 80 30 5 235 102 1.5 1.5 18 547 COMPLETE CAST 305 100 30 5 235 165 19 19 23 684 BIMETAL TIP CAST 305 100 30 5 235 145 90 1.9 1.9 23 684 BIMETAL TIP ASSUME CAST 255 100 30 5 235 105 90 2.5 2.5 30 889 INCOLOY 800HT COMPLETE ASSUME BAR 305 100 32 21 240 435 39 3.9 47 1410 BIMETAL TIP BAR 305 100 32 21 240 270 52 3.9 3.9 47 1404 BIMETAL TIP ASSUME BAR 255 100 32 21 240 180 52 5.1 5.1 61 1825 JKS 109 COMPLETE CAST 305 100 30 26 150 9.1 9 1 EH NICROFER G3 COMPLETE ASSUME BAR 305 100 40 22 240 480 11.7 11.7 140 BIMETAL TIP ASSUME BAR 305 100 40 22 240 300 52 11.7 11.7 140 182 BIMETAL TIP BAR 255 100 40 22 240 200 52 15.2 15.2 ™ INCONEL COMPLETE ASSUME CAST 305 80 55 11 170 145 2.6 2.1 25 760 COMPLETE CAST 305 100 55 11 170 235 3,3 3.3 40 1188 BIMETAL TIP CAST 305 100 56 11 170 205 90 3.3 3.3 40 1188 BIMETAL TIP ASSUME CAST 255 100 55 11 170 140 90 4.3 4.3 51 1544 INCONEL 601 PIPE(40SCHED) PIPE 325 55 60 23 130 180 50 7.0 1.1 13 396 PIPE(80SCHED) ASSUME PIPE 325 73 60 23 130 360 50 7.0 2.2 27 806 Notes. 1. 80mm castings have 14mm walls. Z 100mm castings have 24mm walls. 3. Inconel (40 shed) ppe has 4mm wall. 4. Expected Hfe based on 1993 Portland life of 12 months. 5. Bimetal plunger has a 150mm steel top half. 6. "Relative life for wall O.D." considers the wall thickness and O.D. assuming linear wearrate viz. an 80mm O.D. has 80% of the life of a 100mm O.D. plunger. 7. Cost of maching tip of bimetal plunger is $40 except for Niresist which is $55 8. Assume 255mm CAST STAINLESS materials have 2.5 times longer life than 305mm based on 310SS experience. 9. Assume 255mm BAR has 30% longer life than 305mm based on 304SS experience. 10. Assume 305mm tip and COMPLETE have similar wear rates viz. based on 304SS and Nicrofer experience. CHAPTER 5 PLUNGER WEAR PAGE 215 5.5.4 Plunger Shape In integrated feeder plants, it is common practice to use flat beveled shaped plungers of similar shape to plunger C in Figure 5-2, but most independent feeder plants use flat bottomed plungers. The reason for the tapered shape is the belief that the taper acts like a wedge to crack the hard crust. However, there are several potential benefits of flat bottomed plungers: (i) less bath sticking due to smaller surface area, (ii) larger hole due to "fracturing" rather than "drilling" a hole, (iii) less chance of sticking in the crust, (iv) less chance of heat retention due to smaller surface area touching the liquid bath, so less plunger buildup, (v) less chance of the plunger shaft tracking off centre which causes uneven and premature seal wear, and (vi) higher vertical component of pressure. It is difficult to establish if these benefits have been attained as they are difficult to measure. It is difficult to find a reason to retain the tapered tip of the old design. Comparing wear rates before and after the tip change showed no statistical difference (Table 5-X). TABLE 5 - X EFFECT OF TIP SHAPE ON PLUNGER WEAR RATE NEW OLD CAST STROKE DWELL TIP OVERALL ALL DATA POTS POTS IRON SPEED TIME SHAPE MEDIAN SPEC ONLY ONLY HR Fast 2s Bevelled 5.1 4.5 5.9 4.9 HR Fast 4s Flat 5.6 5.0 6.8 5.4 Statistical y significant (90%)? ^Q!^| Note: All values are wear rates in mm/month for plungers from production pots. CHAPTER 5 PLUNGER WEAR PAGE 216 If there is no wear rate benefits for tapered plungers and there are several operational benefits to flat bottomed plungers, it is concluded a flat bottom is the preferable shape for plungers. T3 changed to flat bottomed plungers in 1988 and Portland changed in December 1990. 5.6 PLUNGER MATERIAL 5.6.1 General Smelter Experience There is a wide range of materials used for plungers in smelters. Table l-II shows what materials some 24 non-Alcoa plants are using for plungers. This is in addition to the nine Alcoa smelters which have used cast iron (until Portland changed to cast 31 OSS in 1993). There are over a dozen different materials used with cast iron being the most common in integrated feeder plants and stainless steel (of various types) is common at independent feeder plants. Several plants have made material changes over time as they have experimented with different types. Plants T4, T5, and T14 have recently changed from cast iron to 304 stainless steel (304SS). T8 changed from 304SS to mild steel (which had an even worse life than cast iron). T4 chose not to go to anything more exotic than 304SS (from cast iron) as they felt that they could not guarantee that the other feeder components would last longer than three years, so why waste money on a plunger that one will probably throw out anyway after three years? This logic is very sensible. It all depends on one's confidence that one knows the life of the other components in the feeder. Good tracking systems and proactive maintenance teams are needed to remove the causes of other failure mechanisms. D3 changed from cast iron to Inconel then to alloy coating and are now looking at other alternatives. Their problem was probably too frequent a break cycle (under 2 minutes) and using alumina as anode cover (which is abrasive). They experienced erosion of the softer alloys. They achieve 2-3 years life, but want to get over 5 years. CHAPTER 5 PLUNGER WEAR PAGE217 In 1987, Pechiney discussed plunger materials at their "Club 180" meeting for users of their API8 (180kA) pots. At this meeting their plants had the following materials (some are trade names). D2 AISI309 D6 Z-12-CN D9 309SS D10 310SS D14 NS 24 D15 310SS D10 are investigating alternative materials for Lines 1 and 2, and are using 310SS for the new Line 3. D6 is changing to a high nickel bar material made by a local manufacturer (Acier Abrasion) with good results. This is the same material used for anode setting crane grabs. The nickel contents of 309SS and 310SS are 12-15% and 19-22% respectively. The fact that the Pechiney plants are moving away from 309SS and 310SS suggest that these are not good enough to match the 5-6 years they are achieving for their crustbreaking cylinders or 7 years for pots. D9 believe they achieved at least three times the life of 309SS when they changed to Nicrofer. The newest line uses Hastelloy X mainly due to economics at the time. (Nicrofer has 40% Ni and Hastelloy X has 45%; both have 22% Cr). The results appear to be at least as good as for Nicrofer. The most recently commissioned Pechiney lines (e.g. D4, D8, D10) have used "310S" stainless steel instead of the standard 310SS materials. The "S" refers to a drop in carbon from 0.2% to 0.08% which is believed to reduce the chance of carbon migrating to grain boundaries and making the nickel less passive to corrosion attack. CHAPTER 5 PLUNGER WEAR PAGE 218 5.6.2 Kaiser Plunger Wear Experiment In 1984, Kaiser published a paper on an experiment into plunger wear (54); this is the only published data traced on this topic. Westerman and Harrison tested a range of cast alloys and steel in a test pot which had three hexagonal plungers. These had strips of the alloys welded around the plunger so a comparison could be made of alternative materials under exactly the same pot conditions. Table 5-XI has the results of the study. TABLE 5-XI KAISER PLUNGER WEAR TESTS ALLOY Ni Cr RELATIVE COST PER COST PER RELATIVE RATIO LIFE TIP TIP-DAY COST/DAY Ni/life HASTELLOY 45 22 23.0 243 0.07 0.18 2.0 309 SS 12 26 6.0 105 0.14 0.35 2.0 304 SS 9 19 4.6 48 0.11 0.28 2.0 13-4 SS 4 13 1.9 - - _ _ CAST IRON 0 1 1.0 30 0.4 1.00 n/a Notes: (i) All materials were cast (not bar). (ii) "Ratio Ni/life" is calculated as follows: (Ni content)/(relative life) e.g. Hastelloy X is calculated as 45/23.0 = 2.0 (iii) All cost values in US$. (iv) Except for "Ratio Ni/life" all data from Westerman and Harrison (5 The conclusions from the paper were that chrome over 20% and high nickel were beneficial to plunger wear. However, they did not notice a more significant trend in their data. There was a linear relationship between nickel content and relative wear rate. The relationship between the life (compared with cast iron) and the nickel content (%) was linear with a correlation coefficient (r2) of 0.99!; 99% of the variability of plunger wear rate was due to nickel rate alone. A 46% nickel content had a life 23 times the life of cast iron, 12% nickel had 6 times the life and 9% nickel had 4.6 times the cast iron life. Clearly, life increases by half the nickel content rise. If one doubles the nickel content, life is doubled. As a corollary, this suggests that one can choose the nickel content of the plunger to provide the desired life. CHAPTER 5 PLUNGER WEAR PAGE 219 If one uses the Kaiser results as a datum and the standard deviation found at Portland (i.e. about 30% of the average (Tables 5-II and 5-IH)), one could use the following logic to determine the optimum plunger nickel content: Target: 95% of plungers last 5 times that of cast iron. Average nickel content: twice the ratio to cast iron. 2x5=10% Variability: Assume that S.D. for the critical ingredient (nickel) is similar to 30% of the average wear rate. Assuming a normal distribution, a 95% confidence limit twice the S.D. below the average would be the lower limit of nickel content. Hence, the target should be: 10+(2x0.6xl0)=22% minimum. Chrome content: Assume Kaiser conclusion of 20% minimum. Hence, final deduced analysis for a plunger 5 times the life of cast iron at Portland would be 22% Ni and 20% Cr (minimum). Portland trials of cast 310SS (20% Ni 25% Cr) gave 3.3 times the life of PA spec cast iron (Table 5-Vin). Extrapolation from 20% to 22% Ni would give an average life of about 4 times the life of cast iron. Hence, this prediction calculation is a reasonable estimate for plant scale operation. Close consideration of the Kaiser report shows that their test may not be completely similar to plant conditions. In the test, the maximum length of service was under 80 days and the dwell time was kept at 6s to accelerate the trial. This longer dwell time tends to make the test one of the ability of the plunger to withstand corrosion only. The effect of erosion and the erosion/corrosion interaction are not tested. Hence, the metals with good corrosion resistance may be seen as the best. Also, the short length of the test may question validity for plant scale performance. CHAPTER 5 PLUNGER WEAR PAGE 220 Based on comments from representatives of several plants which have changed plunger type, the following relationships were evident: Portland cast iron to cast 31 OSS = 3 x original T14 cast iron to 304 = 3 x original D9 309 to Nicrofer = 3 x original T8 steel to 304 = 2-3 x original Thus, Portland trials and plant data tend to support the Kaiser experiment data (if not the Kaiser conclusions). The data suggest that increasing nickel content reduces plunger wear rate, but that a 20% chrome content is also needed. 5.6.3 Portland Experience One can see that the industry is in a continual state of flux with respect to plunger materials. Often these changes are made on limited trial data and the changes are generally on a full scale plant basis. This study considered alternative materials in a controlled manner using statistical techniques and a reliable tracking system. Portland has conducted trials on 19 different materials to date as shown in Table 5-XH Using data from this table, Figure 5-12 illustrates the relative wear rates of different plunger types, and Figure 5-13 shows life based on within-pot trials; same pot, same time, same original size. Below is a discussion on the different materials used for the plungers tested in this investigation. Section 9.3 examines the economics of these materials and designs. 5.6.3:1 Cast Iron In summary, the wear rate of HR is 40% better than PA despite being half the hardness. Based on recommendations from Alcoa USA, cast iron plungers of 21%, 25% and 27% chrome were tested at Portland. Overall, the wear rate of the high chrome cast iron CHAPTER 5 PLUNGER WEAR PAGE 221 TABLE 5 - XH PLUNGER MATERIALS TESTED AT PORTLAND CAST IRON HR SPEC PA SPEC BRITISH ALCAN SPEC 21% CHROME 25% CHROME 27% CHROME ALLOY COATINGS AI2369 AM 721 AH 795 AM 236 AI2354CO STAINLESS STEEL 310 CAST 304 INCONEL PIPE INCONEL CAST INCOLOY 800HT NICROFER G3 JKS SILICON CARBIDE materials was 70% better than PA spec and about 50% harder. Some fracturing occurred from temperature shock or brittleness, but the main problem was plunger buildup (dags). At Portland, it was found that there was almost 3 times more likelihood of having dags on high chrome plungers than any other plunger type (Section 6.3.1). Some 60% of multidag feeders were high chrome, despite only having 13% of these types in operating pots. Due to the hazards experienced with dags, high chrome plungers are not recommended despite the better performance in wear rate. Cast irons were found to be inferior to almost all other materials with respect to wear rate and cost/day (Section 8.3) and are not recommended. CHAPTER 5 PLUNGER WEAR PAGE 222 5.6.3:2 Inconel Inconel was selected for trial as it was a high nickel-based material. There was concern that it may be too soft (half the hardness of cast iron), but it had a proven record at SI on some parts of Soderberg pots. The design used for trial from February 1991 was a 2 inch pipe (60mm O.D., 52mm I.D.) with a plate crimped onto the base. Unfortunately, the plate fell out within 3 months after going into service. This exposed both sides of the pipe to attack and left the outer edge to peel back as it hit the crust. This led to attack of the steel plunger shaft underneath. This is illustrated in Figure 5-2 plungers Fl, F2andF3. Despite these setbacks, the performance was very good. After seven months of operation, the wear rate on one was only 3mm (or 0.4mm/month). Cast iron wear rate was about 7 times that of the Inconel pipe. Wear rate was very low for thefirst yea r until the steel inner shaft was exposed. If one used a thicker walled pipe, the life would be much longer. In order to try a cheaper alternative, a cast version of Inconel was tested. Cast Inconel had a relative wear rate 3.3 times better than PA spec. Although twice the wear rate of pipe, the wear rate was still three times better than PA spec cast iron (Figure 5-11). The cast material had a life about half that of similar analysis bar material. This better life may be due to the quality of the metal grain structure relative to a cast material. 5.6.3:3 Niresist and Incoloy Because of the cost of Inconel and the difficulty of purchasing bar in Australia, a cast material of high nickel was tested. Cast Niresist D4 has a 30%Ni content but only a 5%Cr, so was not strictly what the Kaiser tests recommended. It has considerable carbon content (2.6%) which may migrate to the gain boundaries and may not be attractive for corrosion efficiency. However, it was easily castable and quick to trial and may indicate the effect of a compromise between Inconel and cast iron. CHAPTER 5 PLUNGER WEAR PAGE 223 Ni-resist achieved 1.9 times the life of PA spec cast iron. Although this was significant, this was not high enough to consider it further. The lower life may have been due to the low chrome content. A bar equivalent to Niresist was sourced. Incoloy 800HT had attractive temperature and corrosion statistics, so a sample of 20 plungers was placed in service in 1992. Relative wear rate of Niresist was 1.9 versus 3.9 for Incolov. Incoloy proved to be twice as good as Niresist, once again confirming that bar is superior to cast for a similar chemical analysis. 5.6.3:4 Stainless Steels A material 4 times better than cast iron was found (Incoloy), so a significant breakthrough had been achieved. However, this was still not good enough to get a material where ALL plungers achieve over 5 years. An average life of about 7 years was required. The pattern so far suggested that the high nickel was attractive and that bar gave twice the life of cast materials. Hence, study continued using high nickel stainless steels. Most stainless steel bars are difficult to get in Australia as most are only made overseas and costs per plunger are high. 304SS is the cheapest ($255) and easiest to obtain, with the cost increasing with the nickel content to about $480 for Nicrofer (versus about $40 for PA spec cast iron). Based on successful experience in other smelters, 304SS was tested at Portland with original and 50mm shorter lengths from mid 1993 (Figure 5-2 plunger H). It was later established that the original 305mm length gave a life 5 times that of PA cast iron. At the time of commencing the 304SS trial, it was decided that action should be taken to change normal repair plungers to a high specification design in July 1993 viz. cast 310SS. It was judged that the relatively cheap cost ($120) made it a good economic risk. In addition, following testing on reduced penetration (Section 5.4.2), it was also decided to use 50mm shorter plungers from July 1,1993. This dropped the cost to $80 making the changeover even more economic. CHAPTER 5 PLUNGER WEAR PAGE 224 The effect was a life 8 times that of PA cast iron which achieved not only a better life but cost only 30% that of a 304SS bar plunger. This illustrates the benefit of cheap cast materials (with less penetration) being very cost effective compared to traditional designs using high specification bar material. (Section 9.3 discusses a bimetal tip that halves the cost of bar materials to make them more cost competitive.) After discussions with D9 who had good experience with Nicrofer, a trial was introduced in November 1993 with 50mm shorter plungers. This material was an extension in the trials to date as it had higher Nickel, a good chrome content and a proven plant record. However, no comparison tests had been conducted at D9 and this was an independent feeder plant which may achieve better results due to lower wet time rather than better material. It appeared to be a good risk. It was very difficult to obtain for trial quantities desired, so about $80,000 of material had to be purchased for 400 plungers. To offset the cost, a bimetal design was jointly developed by Hans Kempe and the author to get the maximum number of plungers from the expensive material. Life 15 times that of PA spec cast iron was achieved. It is clearly the best material tested if the target is maximum life. Even though the initial cost is high, it will be seen that the cost/day is still economic (Section 9.3). It is expected that an average life of 15 years can be achieved with Nicrofer. This life may be a little too long, as too long a life may result in disposal of a significant amount of the costly material on failure of other feeder components. That is, if other major components fail and result in the feeder being returned to the workshop, it may not be prudent to return a plunger to the plant knowing that the plunger may be the reason for the next failure in under a 5 year period. Hence, the old plunger will be disposed of, thus throwing away a considerable investment. The challenge of feeder design is to have all parts fail at the same time. Too long a life on some components may in fact be bad economics. If a plunger shaft breaks, for example, the plunger may fall into the pot and cannot be dragged out as it welds itself to the bottom of the pot and does not dissolve. Hence, the pot becomes more unstable. CHAPTER 5 PLUNGER WEAR PAGE 225 To investigate a cheaper alternative to Nicrofer and 310SS, a high grade cast material (30% Ni, 26% Cr, 5% Mo) was invented by the author with assistance of Ken Deans from Backwell/IXL (the suppliers); called JKS...the "Jim Kissane Special". The analysis was based on the highest nickel content that can be easily cast with a minimum of 20%Cr and some Molybdenum as used in Nicrofer and Incoloy 800HT for grain strength. This was introduced in March 1993 on trial feeders. JKS wear rate was 9.1 times better than PA spec. This is comparable to Nicrofer but considerably cheaper to manufacture viz.: $150 versus $300 for a full length cast plunger of JKS versus bimetal tips of Nicrofer respectively. It is also made in Australia rather than fully imported as is the case with (bar) 304SS, 309SS, 310SS and Nicrofer. It is probably the most cost effective plunger design tested to date and is not too long a life (3,000 days) (that the investment does not take excessive periods to achieve a return). Life is also adequate to achieve an average life of about 9 years, which is sufficient so that few plungers will fail under 5-7 years. 5.6.3:5 Silicon Carbide Silicon carbide has been trialed at Portland on two occasions (76). These were rectangular with chiseled sides and a taper at the end for structural support. Ceramics are stronger and easier to make if made rectangular in shape. There is no reason why a plunger needs to be round. In the trial, the plungers fractured half way down their length after a maximum of 2 weeks (Figure 5-2 plunger D). Some lasted less than a day. Analysis of the material showed it was only 60% as strong as the specification. Another trial was carried out where the material was checked at independent laboratories by the manufacturer (Carborundum) and by Portland Aluminium (CETEC) to be above the specification (Figure 5-2 plunger E). CHAPTER 5 PLUNGER WEAR PAGE 226 Once again, the plungers failed within days. This material suffers from fracturing, as silicon carbide is brittle. They break easily if hit sideways. Trials at D10 also failed even with a stroke time of 3.0s compared to the 1.6s at Portland. This material is not recommended. 5.6.3:6 Alloy Coated Cast Iron Four types of alloy coated cast iron were placed in service in parallel with other trial plungers in February 1991 (Figure 5-2 plunger C). None achieved much better results than cast iron (75,76). Some of these alloy coated materials had the coating shed off the plunger within months, possibly due to different values of expansion coefficients. A further trial was conducted on one cast iron plunger coated with Inconel but further trials were not conducted due to the better performance of bar materials. T3 are having some success with a Castolen coating. D3 have decided their alloy coating has not been successful and are searching for a new material. It is the belief of the writer that alloy coatings are a "Band-Aid" treatment and are not thick enough to ensure long life. Hence, coated plungers are not recommended. 5.7 CONCLUSIONS ON PLUNGER WEAR MECHANISM In the foregoing discussion in Section 5.2-5.6, it has been established that there are many possible contributors to the large variability in plunger wear rates (Table 5-1) which cause the large scatter in wear profiles in Figure 5-3. There were two theories on the wear mechanism - erosion and corrosion. A few general observations may be useful to focus on the critical discoveries from the many tests conducted in this study at Portland. The results of this research are summarized in Table 5-XIQ. ui a. o ? £ DC < 5 2 z j" g o CM R z 0- Q_ Q- o o 3 3 3 03 111 8 8 8 z 8 r- E ro O Mi £ nc 2 ui o CD Q. ro o tr a. EC |U < LU i Q z 3 Ji tc 23 o Q g 0_ 0- < 3 a z ui 3 3 o 8 8 8° UJ < LU 2 z z h- «t 0- N (0 2 > £ X 3 1 ® CO o 8 J? ra _1 to ra-" CO < o o o += o 2 cr 2 | o < CO o 3 CO 0 E s ra o to to < to o 5 CT1 J. 5 £ I- o Trr Li. 1— ac 03 1i- 1 LU ti E o c a 2 £ LU III 1- rr o o O i_ a> o z 2 m 3 z a T 1- < rr co O 0. n G ro Q- ^ <0- 0- to a. o0- i- 0. co LU in (fl = f/1 _i Ho 7 Q I d) ra 8 (/) ro 5 rr z _l1 CO CO o t/i X s z o O ra 2 CD U u fl n i- LU I OC LU II" 0C fl n () rr o i- T- CN CO E 8 io iO in no o »= a U LO m LO a, Ul l«s ta CO *± ta <- JS ra CD ® S ra O LU < z I o i- a c tn LT 0- ro UJ LU rr 1- CO CO "• ro-2 V- < CO 3 D in rr n z ncio 7 H Of 3 r- 3 IL U CO Q S S | Ul co LU 8 < _j z n I- a 2 3 to n IY o ro o h- £ OC 5 U Ul LU III CQ It co o a OC < * E U- 1- •L>U IT LU OC LU J. tE^ > LU U. LU OC 3 MC LnU LU i- w < Ii_ LU < oc H z o co °«# ta 5 OC I LU 3 LU 3 3 a. y 55 -) r- LU I3 T 3 (4 5 _l 0. R E o LU 0C LU 0- n o _J "- CD 05 CO 3 0. u 2 - CD \~ O c <£ 1. material (higher Ni/Cr the better) 2. penetration depth (less the better) 3. wet time (less the better) 4. plunger buildup (more the better) 5. plunger speed (slower the better) Material and penetration are much more important than all the others. It was also concluded that the wear mechanism was mainly corrosion, with a lesser effect from erosion and that high nickel alloys with over 20%Cr are the best plunger materials. Bar materials give double the life of the equivalent cast analysis plungers. CHAPTER 5 PLUNGER WEAR PAGE 229 5.8 MAJOR FINDINGS FROM PLUNGER WEAR ANALYSIS (i) Plunger wear was widely variable from pot-to-pot and within a pot. (ii) Wear rates of batch-to-batch cast plungers were statistically in control. (iii) Increasing the frequency of feeding increases wear rate. (iv) Plunger shape (flat bottomed) and mass do not affect wear rate. (v) Wear rate was the same wherever the plunger was located in the pot. (vi) Reduction in speed improved wear rate. (vii) Silicon carbide was brittle and had a very short life. (viii) Optimum material appeared to be a nickel content of over 30% and chrome content over 20%. (ix) Bar material appears to have twice the life of the equivalent cast material. (x) Nicrofer and JKS materials achieved the longest life. 5.9 RECOMMENDATIONS FROM PLUNGER WEAR ANALYSIS (i) Minimize dwell time to that required to release the alumina and control the other factors that cause blockages, rather than operate with an excessive dwell time "just in case". Check occasionally that the dwell time has not been increased. (ii) Investigate alternative materials for plungers only after causes of plunger wear have been addressed. (iii) Control liquid level so as not to get the plunger wet. This may require making the plunger shorter so there is less (or no) penetration in the liquid. (iv) Measure plungers into the workshop to monitor plunger wear. (v) Tracking system should include pot number and location in the pot to tie in pot conditions with feeder life. (vi) Compare plunger materials/designs within the same pot for the same period of time starting at similar sizes to measure relative wear rates (vii) Install Nicrofer bar or cast JKS plungers for optimum plunger life. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 230 CHAPTER 6 PLUNGER BUILDUP ANALYSIS 6.1 SUMMARY In October/November 1993, Portland conducted a feeder survey in order to determine the cause of dagging (buildup on plungers) for 1,000 feeders. It was found that high plunger temperature was the main cause, as well as high chrome cast iron plunger material. High chrome stainless steels bar plungers had minimal buildup problems. Dagging was promoted by high liquid levels and proximity of the flame to the plunger. There was no relationship between frequency of dagging and anode effects. Removal of dags had no effect on the frequency of dags returning. Actions recommended to reduce dagging included no further use of high chrome cast iron plungers, trials of shorter plungers, remove dags only if the dag fouls the feed chute under the superstructure and remove the cause of dagging, not work on the effect. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 231 6.2 INTRODUCTION Buildup of material on plungers is called a range of names which include a dag, dough ball, hornet's nest or ball (Figure 6-1). FIGURE 6 -1 DAGGED PLUNGER It is hazardous for operators to remove dags as they have to lean into the pot with a large crow bar to knock it off the plunger. This exposes operators to dust, heat and fumes and they are at greater risk of being splashed by molten bath. The action of knocking off the dag can lead to back injuries due to the awkward procedure. The dags often return after removal, making controlling dags a frustrating, as well as a hazardous, job. Dagging can suddenly become frequent on many pots then reduce again after a few days. Operators note that dagging seldom occurs on all feeders in a pot, except for new pots (less than 30 days old) when dagging often occurs across the pot and the dags are very big (>400mm diameter). Pot operators also observe that the feeders in the middle of the pot dag more frequently. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 232 For integrated feeders, operators must remove the dags if they affect the plunger from fully retracting. The dosing unit on integrated feeders is activated by the same cylinder that moves the plunger. If the plunger does not fully retract, the dosing unit may remain open causing alumina to flow continuously unmetered into the pot making it mucky. Conversely, if it does not retract to its top limit, the dosing unit may not refill at all, causing the pot not to be fed and anode effects will result. Usually the dosing unit will be affected if the dag fouls the chute under the superstructure. Independent feeders are not affected by dags except if the dag fouls the chute in such a way that the alumina dose is deflected to the side of the hole in the crust. This is rarely the case. Dosing accuracy is not affected by dags in independent feeders as the crustbreaker is separate from the dosing unit. On Pechiney and most other independent feeders, there is a dag cleaner for continual cleaning of the plunger. Refer to Figure 1-12 and 1-13 which show the dag cleaner at the base of the feeder around the plunger. These cleaners are relatively successful if the cylinder is 150mm O.D. or bigger. Below this diameter, there is insufficient force to fully retract the plunger if there is buildup. Even for the 200mm cylinders used at D9, the plungers still get stuck if the pot is new or the bath level is high (which both cause dags). For many pots with smaller cylinders, one can often see trials of dag scrapers which have been unsuccessful. Many trials were conducted at Portland before the dag cleaners were raised to above the normal high bath level. Since then, there has been no problem with dags and feeder faults. The real solution is to stop the creation of dags rather than use questionable designs to correct for an operating fault. Possible causes of dagging are: (i) the plunger is too hot to allow the liquid bath to freeze and fall off, (ii) the bath chemical composition is low in aluminium fluoride causing material to freeze, and (iii) the surface of the plunger is rough causing the material to "key" into the surface perturbations. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 233 Some people believe that dags cause anode effects due to restriction of flow of alumina into the hole in the crust. An investigation was conducted to determine what causes dags and develop ways of preventing them from occurring. A survey on 200 pots was conducted during October/November, 1993 to ascertain which feeders had dags and what relevant pot conditions existed at the time. Pot attendants logged, every time a dag occurred, the (i) pot and feeder location of any dag, (ii) size of the dag, and (iii) if the dag was removed. Table 6-1 has a summary of the observations. As it was clear that there was a relationship to the temperature of plungers, plunger temperatures were measured for a range of different pot conditions in a separate trial. This Chapter firstly discusses the dag survey results, explains the investigation into plunger temperatures, then discusses how to stop dags from occurring. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 234 TABLE 6 -1 DAGGED FEEDER SURVEY SUMMARY (29 Oct 1993 - 12 Nov 1993) NUMBER OF POTS SURVEYED 204 (1040 feeders) NUMBER OF DAGS 43/day (4% of feeders) BIG DAGS (>400mm) 26% NUMBER OF DAGS REMOVED 10% FEEDERS WITH MULTIPLE DAGS 3% of feeders (34% of total dags) POTS WITH >1 DAG/DAY 4% of feeders (25% of total dags) POTS WITH NO DAGS EVER 45% FEEDERS WITH NO DAGS EVER 81% DAO DISTRIBUTION BY LOCATION OF FEEDER IN POT DISTRIBUTION OF DAQS BY PLUNGER TYPE DISTRIBUTION OF PLUNGER TYPE HIGH CHROME 11% HIGH CHROME S0% POTS THAT HAO ANY DAGS FEEDERS THAT HAD ANY DAGS DURING OBSERVATION PERIOD DURING OBSERVATION PERIOD NO DAGS EVER «% NO DAOS EVER 91% CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 235 6.3 SURVEY RESULTS 6.3.1 Dag Frequency Dags occurred on 4% of feeders per day but 81% of the feeders did not have one dag at any one time in the two week period. Of the pots that had dags, only 4% had more than one on a pot in any one day. Operators suggested the frequency of dagging was about what is commonly seen. Dags tended to come and go apparently randomly irrespective of whether they were removed or not, except for the high chrome plungers which had higher than expected frequency of dags. 6.3.2 Plunger Type 30% of all dags and 60% of repeaters were on feeders with trial high chrome cast iron plungers despite these only comprising 13% of feeders in the plant. There was no trend with any other type of plunger in the initial survey. Following concerns of dagging on trial 31 OSS bar plungers at T9, a further survey was conducted in March, 1994. Portland was not using 31 OSS bar, but some Nicrofer trial plungers also contained high chrome and high nickel which one would have expected to behave similarly to 310SS. It showed there was 16% less chance of dagging on Nicrofer plungers than for normal cast iron plungers. One observation which is common for cast materials is the rough surface on some plungers. This may provide a place for bath to freeze and "key". These may dag when first installed, but eventually the surface will become smooth. The bar material starts off very smooth and is not affected by manufacture. In several plants, trials of different cast materials have been unsuccessful despite the plungers working satisfactorily at other plants e.g. 31 OSS at T6 and PA spec cast iron at T9. This may be due to poor manufacture quality control rather than a fault with the material per se. 6.3.3 Dags Removed One operator removed all dags during the test period, but the number of dags was the same in his pots over the period as the other pots; removal had no effect. Very few dags CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 236 were removed by operators (10%). Only extremely large ones (>400mm diameter) were removed as these were fouling the feed chute. 6.3.4 Anode Effects No trend was seen on anode effects (Table 6-II(a)). In fact, there was more than double the chance that no anode effects occurred if a dag was observed viz. 10% versus 4%. This suggests that dags cause the holes to be larger, so there is less chance the hole will block and restrict alumina flow. It may also suggest more alumina bypass unmetered into the hole on these integrated feeders which results in overfeeding and potentially more chance of muck forming. Hence, dags are not significantly linked to anode effects. 6.3.5 Liquid Levels Table 6-II(b) shows the percentage of pots (in a given total liquid level range) that have dags. "In spec" pots are those which the total liquid level is "in specification". The "Dag Factor" is the number of dags divided by the number of pots (in that liquid level range) divided by a similar calculation for the "in spec" pots. It is a way of quantifying the chance of dagging by showing the relative frequency of dagging compared to ideal pots. For example, if the total liquid level is 40mm above normal, there is almost 3 times more chance of dags, than if the total liquid level was below normal. Hence, there was more likelihood of dags (especially multiple dags across the pot) for high liquid levels. 6.3.6 Location in Pot The middle of the pot dagged the most with 32% of dags occurring on the middle of five feeder locations (Table 6-1). This trend is similar to plunger temperatures, suggesting a relationship between dagging and plunger temperature (Figure 6-2). If bath chemical concentration was important, one would have expected that there would be similar frequency on all plungers and that all plungers would dag at the same time. This was not the case, so this discounts the theory that bath chemistry is the cause of dagging. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 237 TABLE 6-H LINK BETWEEN DAGS AND ANODE EFFECTS (a) DAGS FOR DIFFERENT ANODE EFFECT HISTORY NUMBER OF DAGS NUMBER OF ANODE EFFECTS ON ONE DAY ON ONE DAY 0 1 2 3 4 5 0 1020 296 56 15 8 3 1 115 36 8 1 2 33 10 1 3 18 4 4 2 3 5 3 TOTAL DAGS = 1632 NUMBER PERCENT ANY AE OCCURRED? ANY AE OCCURRED? YES NO YES NO DAG? YES 63 171 4 10 NO 378 1020 23 63 (b) PERCENTAGE CHANCE OF DAGGING RELATIVE TO TOTAL LIQUID LEVEL LIQUID LEVEL #POTS NUMBER OF DAGS AT EACH FEEDER LOCATION (%) DAG RANGE IN RANGE 0 1 2 3 4 5 FACTOR NORMAL 51 65 24 10 1 1.0 >20mm 71 54 20 15 7 4 1.3 >40mm 51 52 13 9 6 6 14 2.8 Notes: (a) "Normal" liquid level is the total (bath plus metal) liquid level for normal pot operation. ">20mm" and ">40mm" are levels 20mm and 40mm in excess of normal operating total liquid level. (b) The "dag factor" indicates the relative tendeny to dag. It is calculated as follows: ((# daqs/# pots) in chosen liquid level range) ((# dags/W pots) in normal liquid level range) CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 238 FIGURE 6 - 2 PLUNGER TEMPERATURES FOR 945°C POTS BY LOCATION LOCATION 1 MEDIAN 11i"C Ott 3 -t-H—I—I—I I I I I 91-100 131-140 171-180 211-220 261-280 291-300 331-340 LOCATION 2 MEDIAN 1S6°C 81-100 131-140 171-180 211-220 261-260 291400 331440 LOCATION 3 MEDIAN 168% 91-100 131-140 171-180 211-220 251-280 281400 331440 LOCATION 4 MEDIAN 146"C 61- 71- 91- 111- 141- 161- 171- 191- 211- 231- 261- 271- 291- 311- 331- 60 80 100 120 140 180 180 200 220 240 260 280 300 320 340 LOCATIONS MEDIAN K6°C 61- 71- 91- 111- 131- 161- 171- 191- 211- 231- 261- 271- 291- 311- 331- 60 80 100 120 140 160 180 200 220 240 280 280 300 320 340 Motes: (I) Plunger temperatures (°C) for 8 pots on 20.8.93, 31.8.93 and 10.9.93 (ii) Overall plunger temperature: mean 143°C, median 135°C and s.d. 42°C. (iii) Pot temperature: mean 945°G and standard deviation 17°C (iv) Y axis represents number of observations. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 239 6.4 CONCLUSION ON CAUSES OF DAGS From section 6.3, high chrome material was clearly the most important factor in dagging. However, this material was used only on trial and is an exception. This trial aside, it appears that dags are caused by (i) surface roughness, (ii) high bath level, and (iii) high plunger temperature. Surface roughness can be addressed by use of bar material or by quality control of cast manufacture. For normal operation of a pot, there is little that can be done to do anything about the plunger on operating feeders. High liquid levels would tend to have material contact at higher levels on the plunger which are not as smooth as the normal operating surface of the plunger by the increased wetted surface area. This may give a starting point for gradual buildup as the bath freezes. Control of bath levels is via control of anode cover bath/alumina blend, tapping/filling pots if out of bath level target range, cover practice when setting new carbon and cover practice over holes during normal operation. High liquid levels not only have an affect by coverage of higher surfaces that may be rough, but also affect the surface temperature of the plunger. There are other causes of high plunger surface temperature such as flame from the hole, radiant heat and high frequency of crust break cycles. It is postulated that high plunger temperatures cause dags by the interaction of late freezing and loose powder around the plunger. Because the plunger is hot, the bath temperature is much higher than its freezing point (usually about 920°C) and any particles of alumina or anode cover in the area will tend to cling to the wet bath. This causes surface roughness and further plunging into the bath allows continued buildup to occur until eventually, if the dag gets big enough, it may fall off due to hitting the feeder chute, dag scraper (if installed) or the crust. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 240 The A2 feeder showed quite clearly the link between buildup and this procedure. The A2 plunger raises and then the alumina is delivered. Because the plunger is often still wet, some of the alumina will stick on the surface and the dag is created. Once the dag starts to form, the rough surface allows buildup to continue by "keying" onto the surface. The next section discusses studies on the causes of high plunger temperatures so that a better understanding of one of the main causes of dagging can lead to methods of dag prevention. 6.5 FACTORS AFFECTING PLUNGER TEMPERATURES For a range of pot and feeder conditions, 120 measurements were taken of gas temperature next to plungers when fully retracted. The results are discussed in the following sections. 6.5.1 Pot Temperature The spread of plunger temperatures was very wide and was independent of pot temperature (Table 5-VH). The higher the pot temperature, the higher the plunger temperature... 10-20°C rise in plunger temperature for similar rise in pot temperature. However, the variability was extremely large. The peak plunger gas temperatures were independent of pot temperature. 6.5.2 Plunger Temperature Over Time The scatter of temperatures of a 3 week period also showed temperatures varied over time by a large margin (Table 6-ID). It should also be noted that pot temperatures averaged 945°C for this survey which is very low for pot operation. Normal pot temperatures are generally hotter than this - about 955-960°C Despite this, the peak plunger temperature reached over 300°C (versus a mean temperature of 143 C). There was a random variation in temperature for the 40 plungers, with large swings in plunger temperature independent of pot temperature e.g. 1007#3 dropped 180 C for a rise of 35°C in pot temperature and 2074#3 increased 85°C for a drop of 2°C in pot temperature. fS o < OH CO CD 0 CD E E E CU cu CD m CO CO CO CL a. CL LL LL LL O O O o 111 —. o o o o O o o a o o O o O m o i m o i CO o • a o CO • in in o O o O *- CM T- CM T- CM 8 CO T- CN in n a: r> U0 J •-. y CO CO "^ CM o> CO r^ r~ in o CN in CM o "3" Z CL. CO c cu cu CU CZ 0 a> E E E CD E m a. CO CO CO a. O LL LL LL O LraL o ui —« O O o o O O O -• E o a • UJ y '^• CO CM •* r- co ^- N- in CO o OO CO CO CO N -^ CO a in •^r o CO r^ in zv Qr- CM CO CO T -«- co r* CM "3- CO in CO ,C—N CO o co rr in vr o in •a- OS ^' CM -1 Ul Q- 1- CO CD a> cu c cu cu c cu c 0 E E E CD E E S E CD E in co _. o m CO CO Q. (0 CO J9- co CL co -^ Q! CM LroL LL LL O LL LL LraL z O on o 1- o UJ? o o O O o o O O o o o o o o CO CM o i CM O i f- o ' CO o o 8 i o a o o o o ZU9 5 •>- T •<- CM T- m CM t- T CM T— CO 1 CN o o CO *~ _BCl UJ Q K§ LU ftp UJ U " in co o o> m s •a- ai co IO T- O) T OIOS T- CO o CM CO TT o LL Z Q. CM CM r> M2 CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 242 Figure 6-2 shows the same data sorted by location in the pot. There was an overall trend of #3 being the hottest, followed by 2 & 4 (10-20°C lower), then 1 & 5 (a further 20-3 0°C lower). The middle of the pot had the highest temperature. This was similar to the pattern for dags discussed in Section 6.3.6 and illustrated in Table 6-1. The reason for the highest temperature in the centre of the pot is probably due to the pots being side-by-side (similar to the layout in Figure l-4(b)) where there is less air circulation in the space between pots. This the normal layout for prebake pots. End-to- end pots (such as thefirst three pots on the left hand side of Figure l-4(a)) may not be affected as much by this problem as there is better air circulation along the complete length of the pot. 6.5.3 Hole Size and Flame Table 6-III also shows the hole dimensions. If the hole had a strong flame that is near or touching the plunger, the temperature reached up to 300°C If the hole is closed, plunger temperatures were about 100 C This variability was far greater than variation from location in the pot. The size of the hole interacts with the amount of gas flow at the plunger. The stronger the flame under the plunger, the higher the plunger temperature. If the hole is large, the gas can escape more easily, so the plunger is generally cooler. 6.5.4 Number of Holes Open The number of open holes affects gas velocity at the plunger. If all holes are open or if all holes are closed, plunger temperatures are cool. If only one or two holes are open, then the high gas velocity tends to impact the exposed plunger more, raising plunger temperature. A survey of 44 random pots (220 feeders) at Portland showed almost half the holes were closed (Table 6-IV). In comparison, T17 rarely has holes open, and Til has CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 243 under 30% open. Both plants have 3 feeders per pot and little to no plunger penetration. D9 generally has all of their 4 feeder holes open, as they have deep plunger penetration. Also, operators check there are no blocked feeder holes every 32 hours as part of their pot checks. TABLE 6-TV SURVEY OF FEEDER HOLES OPEN ON OPERATING POTS #HOLES OPEN POTS WITH HOLES OPEN NUMBER PERCENT 0 21 48 1 12 27 2 9 21 3 1 2 4 1 2 5 0 0 Note: Survey of 44 pots (220 feeders) at Portland. It is best to either have all holes open or all holes closed, then gas velocity is minimal at the plunger. Venting the gas elsewhere from the feeder or low penetration will help prevent dags (but will have a negative effect on heat loss). 6.5.5 Plunger Length and Penetration In the survey of 44 pots at Portland, a number of feeders were test pots where the plunger length had been shortened by 50mm as part of the trials on optimizing penetration depth (Section 5.4.2). Short plungers had a 40% greater chance of having the hole closed compared to the historical length. This is due to less bath penetration with short plungers and more chance for anode cover to fall into the hole to protect the plunger from heat. This suggests that plunger temperatures are less and there will be less dagging for the shorter plungers. Operating pots at a target total liquid level (rather than just bath level) optimizes feeder life and dagging. The lower the penetration, the lower the plunger temperature and the better the wear rate. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 244 6.5.6 Proximity of Plunger to Hole Table 6-V shows temperature at different distances from the chute. This indicates the effect of proximity of the plunger to the hole. The closer the plunger was to the anode cover the hotter the plunger. Every 100mm closer raises plunger temperatures by about 50°C This has been confirmed by similar tests by Patrick at Boyne Island smelter (83). The higher the anode cover or the closer the anodes are to the plunger the more the hot jet of gas impacts the plunger and the hotter the plunger becomes. Deep anode cover does not help dags. The deeper the metal, the higher the anodes and the hotter the plunger. Once again, liquid level affects dags. It is useful to explain the relationship between crust level and liquid level. The pot voltage is controlled by moving the bridge, and hence the anodes, up and down to maintain a target resistance between the bottom of the anodes and the top of the cathode (Figure 1-3 and 1-5). There is generally a gap of 40-60mm between these two surfaces. As the metal rises due to metal production, the anodes are raised and the crust attached to the anodes rises as well. Thus, the hole gets closer to the plunger until the metal is tapped and the metal level drops again. Smelters tap metal usually between 24 and 48 hours apart. Clearly, the plants with longer times between tapping expose plungers to more period of high temperature and have relatively greater chance of dagging. If a pot becomes unstable, the resistance often rises to move the anodes away from the unstable metal level. Once again, this causes the cruot to approach the plunger and exposes the plunger to higher temperatures. Note also that bath level rise does not cause the crust to rise, as the anode height is purely affected by the anode-cathode distance. However, the higher the bath level, the deeper the plunger will be wetted, so dags can form on the surface. The ideal is to have a large feeder stroke length to keep the plunger as far away from the hole as is possible. It is interesting that D7 actually raise/lower their crustbreakers in the superstructure to achieve minimum bath penetration. In this way, they effectively CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE245 TABLE 6-V PLUNGER TEMPERATURES RELATIVE TO CHUTE TEMPERATURE ACTUAL TEMPERATURES (°C) TEMPERATURE RISE (*C) DISTANCE BELOW CHUTE (mm L_ DISTANCE BELOW CHUTE (mm) TIP TEMP (*C) ZERO 100 200 300 400 100 200 300 400 <100 98 107 107 107 100 9 9 9 2 90 106 - 16 - 75 87 95 12 20 92 122 167 240 - 30 75 138 - 79 150 262 •QUI - 71 183 430 - 96 106 107 10 11 100-150 126 211 - 85 184 - 105 120 231 - 14 126 - 136 220 tfbfm - 84 267 - 145 212 - 67 - 148 250 KH - 102 242 - 110 105 105 100 117 5 5 7 >150 350 360 442 532 +10 92 182 233 280 443 578 780 47 210 345 547 217 279 344 430 476 | 62 127 213 259 Note: Red number s are over 300UC. RELATIVE TEMPERATURE INCREASE BELOW CHUTE (°C) DISTANCE (mm) 100 200 300 MEDIAN 14 85 140 MEAN 35 92 192 STD DEV 33 80 140 NUMBER 9 15 8 MEDIAN AIR TEMPERATURE RISE VERSUS DISTANCE BELOW PLUNGER TIP CO S ui K Ul CL s 150 200 DISTANCE BELOW PLUNGER TIP (mm) CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 246 have a different stroke length, but still have a standard feeder. Hence, dagging and feeder life are optimized. However, this feature is a very costly design and cannot be easily incorporated in most plants. It is preferable to control the cause (liquid level) rather than control the effect (feeder heights). Another factor worth mentioning is that this temperature rise from proximity of the crust can have a very significant effect on cylinder seal life. If the seals are operating just under their maximum rated temperature (as is the case at Portland), a change in distance from the hole could push the seals over their limits and cause failure (Section 7.3.3:3 and 8.4.2). Integrated feeders have a crustbreaker stroke distance of 356mm (except for T8 at 470mm, T8 at 470mm, and Portland and T2 at 508mm). Independent feeders have stroke distances of 500-600mm (Table l-H) so are farther from the crust and hence plunger temperatures are lower. This will assist prevention of dags forming. 6.5.7 Wet Time Although the time that the plunger was wet was not varied in the dag survey pots, trials of a group of 17 pots at different dwell times showed that changing from 4.5s to 2.5s dropped dags by 70%. Wet time is governed by: (a) whether the plunger immerses in the bath, (b) time in the bath per crustbreak, and (c) frequency of crustbreak cycles. Increasing dwell time may have no effect on wet time if the plunger does not touch the bath. Hence, addressing theoretical penetration depth can allow higher dwell times without affecting dagging or feeder life. With respect to penetration depth, Section 5.4.3 discusses plunger wear and Section 8.4.2 discuss the heat conduction up the shaft to the seals. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 247 Generally, dwell time is seldom changed in a plant. Hence, feeder speed, air leaks and valve quality need to be controlled to minimize dags. To test this, thermocouples were placed on a feeder at the top of the plunger shaft and halfway down its length (Section 8.4.2). The temperature trace of the two thermocouples is shown in Figure 6-3. Figure 6-3(a) and 6-2(b) shows the situation where the plunger temperature was initially cool (70°C); (a) has thefirst 60 minutes and (b) has thefirst 200 minutes. Figure 6-3(c) was at ambient temperature when first placed in the pot. Within 16 hours, the plunger and shaft hadrisen t o 250°C despite the fact that the cylinder had not cycled into the bath at any time. This test showed that the temperaturerise is slow (5 minutes to rise 50 C), but within hours the plunger temperature reaches a steady state temperature irrespective of whether it is cycling in the pot or not. Thus, penetration may not be a significant factor in heating plungers, only the effect of gas flow and/or radiant heating. If the plunger is hot, heat will conduct up the shaft to the feeder components. Hence, the whole feeder reaches the plunger temperature within 16 hours. If a dag occurs and if the primary cause is high temperature, the dag will probably continue to reform if knocked off. This theory is substantiated by the observation in the plant dag survey that removal had no effect on dags returning (Section 6.3.3). Whatever temperature the plunger is subjected to, the plunger temperature will quickly report as the same temperature at the cylinder seals. As seals and stainless steel springs cannot take temperatures much over 200 C, the situations where there is a flame at the plunger will rapidly affect cylinder life (Section 8.4.2). This normally results in air leaks, ore leaks, and spring failure. A good rule of thumb is that a feeder with a dag is probably also cooking the seals. Plunger wear may in fact reduce with dags. The buildup reduces the frequency that the outer microlayer of corrosion products falls off the plunger at each immersion (Section 5.4.7). CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 248 FIGURE 6 - 3 PLUNGER TEMPERATURE PROFILES (a) INITIALLY COOL PLUNGER (3 HOUR WARM-UP) 500 •—•" O 400 •— * 0 UJ PISTON ROD,^—•""""""^ .-.Af". - - jjj 300 -j—j_ZJ^j^-z '\>-^r H a 200 0. ^S^ ^ PIUNGER SHAFT Sf 100 1 1 1 1 1 1 10 20 30 40 50 60 70 MINUTES OF PENETRATION IN BATH (b) INITIALLY COOL PLUNGER (3 HOUR WARM-UP) 500 -p.— K PLUNGER WITHDRAWN PISTON ROD t 4 f - O 400 *^>. FROM BATH o • q UJ § 300 — i- T —^PLUNGER SHAFT £ 200 • Q. BJ 100 1 1 1 1 50 100 150 200 250 MINUTES OF PENETRATION IN BATH (c) INITIALLY HOT PLUNGER (12 HOUR WARM-UP) 400 O 300 a Ul BC J- 200 • 111 a. 100 - 10 20 30 40 50 60 MINUTES OF PENETRATION IN BATH CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 249 6.6 DAG PREVENTION AND CONTROL From the previous discussion, it appears that excessive plunger temperature is the principal cause of dagging. Methods to minimize dagging include both prevention and control. 6.6.1 Plunger Coolers One solution to dags is to cool the plunger with air. In 1993, trials were conducted on 6 startup pots in which airlines were installed to cool the plungers. This proved very successful in stopping dags and proved quite clearly that cool plungers do not dag. However, at about $1,000 per pot (plus the cost of air) the solution was too expensive for hundreds of pots. A cheaper mobile cooler was developed which can be moved from pot to pot so only the feeders with problems are treated. Once again, air usage would be high if many units were in use and it would be difficult to control the number in operation as operators would tend to install them wherever a dag occurred rather than control the cause. Coolers would be used to control the effects without addressing the principal cause. This method of preventing dags is not recommended because of capital and operating costs. However, it is a useful tool in the event of a major dagging problem on a very hot pot. 6.6.2 Liquid Level Control As noted in Section 6.5.6, the crust and hole moves farther from the fixed retracted plunger if the metal level falls. The bath level has no effect on penetration through the crust, only on the depth the plunger gets wet. A feeder design must be chosen so that the plunger penetrates every day for all pots irrespective of normal variability in liquid level. It is commonly thought that total liquid level (bath plus metal level) can be controlled by good control of bath levels (as metal level varies little from day-to-day in most pots; CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 250 about 15mm to 30mm). One must also recall that there are different metal levels in each pot. Pots may differ in metal level by up to 100mm across all pots of a plant. As in almost all plants the feeder position isfixed, the penetration depth will vary from pot-to-pot (Table 5-VI). Thus, metal level in each individual pot is important. In addition to pot-to-pot variability, accuracy of tapping will affect total liquid level. Variability in anode cover quality causes a changing addition rate of bath material into the pot so bath levels vary as a result. Good control of bath level is very difficult to achieve in most operating pots. In a pot with a large carbon area (such as Portland) this is even more difficult as a small addition of material may result in a large change in liquid level. Pechiney have invented a bath dosing unit for their API8 pot which is similar to a pot feeder (53). This an expensive but an effective way to ensure a fixed liquid level in the pot. A total liquid level target for each pot would address the metal level variation. These targets should be based on a laser measurement of "tide mark" on the plungers in the pot, such that there is a standard penetration depth for target liquid levels. If the metal reference is changed, then the target is changed. In this way feeders will be standard, but pots will be different. As there would not be one target value for all bath levels, "variation from target" would be used to report on performance of those who remove metal from the pot (called tappers). This would help both dagging and feeder life. The main problems with this method is the variation in bath volume from pot-to-pot. This affects dissolution rates of alumina and aluminium fluoride additions, so could affect fluoride concentration in the pot. High bath levels usually result in the pot being more unstable and less efficient. For pots with high metal levels, the bath level could be quite low, which could affect alumina dissolution and setting. In addition, having different targets for each pot is more confusing to operators, and could lead to less control of liquid levels. As pot control is more important than feeder life, targetting total liquid level is not recommended. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 251 6.6.3 Plunger Penetration Depth Reducing plunger length (without otherwise changing the feeder) will raise the bottom edge of the plunger farther from the crust. This will wet the plunger less, and reduce the gas heating effect. The main problem with this type of action is that some pots may have low metal level (and hence low total liquid level), so may suffer from insufficient penetration. Unfortunately, to cater for the exception, all feeders will suffer from either higher failures and dags (if metal level is high), or will be more prone to blocked feeder holes (if metal level is low). A method has been designed by the author to modify feeders on pots with high metal levels, so that the net penetration will be less. These special feeders will have shorter plungers. A series of pins on the mounting flange prevents these feeders being installed on incorrect pots, and vice versa. It is therefore possible to change (say) 90% of feeders to very low penetration, yet cater for the remaining 10% with special feeders. In this way, it is not necessary to accept poor plunger wear, low feeder life and poor penetration for all feeders in order to protect a few. All that is required is to keep some special feeders as spares for a few pots. Cost is only about $50/feeder. Section 5.4.2 discusses plunger penetration in detail. 6.6.4 Pot Gas Venting It is desirable that the gas from the pot does not vent near a feeder. If the feeder hole is buried (by use of anode cover), the plunger cools and the pot tends to find another place to vent. As long as the real cause of dagging is removed (e.g. anode effects, high bath level, slow feeders) this procedure would be successful. Section 3.3(b) noted that high gas flow can cause blockages in the crust. Thus, high gas flow affects feeder dagging and dose delivery. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 252 Breaking a vent hole away from the feeders reduces the gas velocity at the plungers. This is not possible at many plants as there are no means to break a hole. Breaking the hole at the tap hole (where the metal is removed from the pot) is very difficult to keep consistently open as this is commonly covered to reduce airbura of the end anodes in the pot. Hence, if possible, make gas holes to vent gas away from the feeders. 6.6.5 Reduced Anode Cover Depth Having the anode cover far away from the plunger causes heat loss from the hole and probably burning of the carbon due to exposure to air. Covering anodes is more important to the economics of operating a potline than feeder life and dags. HiaV> carbon consumption and poor pot efficiency are many times more expensive than feeder life. There are other things that can be done to minimize dagging that do not affect pot performance. It is reasonable to comment, however, that increasing the cover depth (as is desirable for minimum airburn and minimum carbon consumption) will deteriorate feeder life and increase dags. It is wise to design and operate the feeder to maximize feeder life without dictating how much anode cover is placed on the anodes. CHAPTER 6 PLUNGER BUILDUP ANALYSIS PAGE 253 6.7 MAJOR FINDINGS FROM PLUNGER BUILDUP ANALYSIS (i) Dags form on hot plungers. (ii) High chrome cast iron plungers dag more than HR or PA spec cast iron or stainless steel plungers. (iii) Dags appear and disappear randomly over time. (iv) Removing normal sized dags has no effect on the frequency of a dag reforming. (v) Anode effects and dags are not related, except for very big dags that may foul the chute sufficiently to freeze the feeder in the retracted position. (vi) Operators often blame dags for anode effects because they cannot think of any other reason. Dags are not a major contributor to anode effects. (vii) High total liquid level promotes dagging. (viii) Proximity of the plunger to flame or the anode cover promotes dagging. (ix) Wet time affects dagging. (x) Rough surface castings can encourage dagging. 6.8 RECOMMENDATIONS FROM PLUNGER BUILDUP ANALYSIS (i) Do not use high chrome cast iron plungers. (ii) Ensure that cast plungers adhere to a surface roughness specification on delivery from the supplier. (iii) Conduct trials on shorter plungers. (iv) Only remove dags that foul chutes butfix th e real cause of dagging. (v) Use longer plungers for sustained low level liquid pots. (vi) Do not target total liquid level for each pot as this may result in lower efficiencey on some pots which may be less cost effective than the savings in feeder life. (vii) If possible, vent gas away from feeder plungers. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 254 CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION 7.1 SUMMARY Joint high frequency trials by Portland and Parker Hannifin (Australia) on 12 cylinders found that the life of seals was affected mostly by (i) temperature, (ii) speed and (iii) eccentricity...in order of priority with the most important first. A cylinder lubricant investigation identified that unless the correct type and correct amount is used, seal life will deteriorate due to lubricant breakdown. Teflon has a lower coefficient of friction and higher temperature limit than Viton. Improved seals and alignment are expected to extend rod and piston seal life and reduce air consumption. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 255 7.2 INTRODUCTION As can be seen from the previous discussion, cylinder design is critical to the life of feeders. Smelters have tried many different cylinder designs over the last 50 years but one is still faced with cylinders that fail within a few hours to over 7 years; even in the one plant. The causes are often difficult to trace due to a poor or non-existent tracking system and lack of data on plant conditions relevant to cylinder life. Clearly cylinder performance has a significant effect on shot size accuracy (Section 2), crust breakage (Section 6) and frequency of failures. It is also clear from Figure 6-3 in Section 6.5.7 that cylinders are exposed to similar temperatures to plungers due to conduction of heat up the shaft, so the pot conditions have a most significant effect on cylinder life. One needs to ensure any comparison of cylinder type or cylinder component type are conducted under cell-like conditions. No papers have been able to be traced that have trialed different cylinder types under controlled conditions comparable to aluminium cell conditions. Cylinder companies quote vague data of how well their products operate on various plants but it has been found that these data are not objective. The author has not as yet heard a voluntary statement from a supplier that their product was no good. In fact, there have been several instances when the author (and other plant representatives) have been deliberately mislead by cylinder suppliers on the assumption that plants would not be contacting each other due to competition. This was quashed when many plants started to converse as part of this research. A number of suppliers found that their products lost favour when the facts emerged. In order to understand the features that both improve cylinder life and the effects of plant conditions on cylinders, two investigations were conducted. Thefirst was a rapid cycling of cylinders of different component under controlled conditions. To date it has not been possible to do trials of different designs, but equipment is available as a result of this research to do this if required. The second investigation was concerned with lubrication of cylinders as this is a major contributor to seal life which is in turn the most frequent part that fails in cylinders. These investigations are discussed in Sections 7.3 and 7.4 respectively. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 256 7.3 CYLINDER TRIALS NOTE: DATA FROM THE PORTLAND ALUMINIUM AND PARKER HANNIFIN (AUSTRALIA) CYLINDER TRIALS ARE CONTROLLED BY A CONFIDENTIALrrY AGREEMENT AND CANNOT BE PUBLISHED OR REPRODUCED OUTSIDE ALCOA OR PARKER HANNIFIN WITHOUT EXPRESS WRITTEN PERMISSION FROM PORTLAND ALUMINIUM AND PARKER HANNIFIN (AUSTRALIA). 7.3.1 Introduction From March 30 to September 3,1992 a joint investigation was carried out by Portland Aluminium and Parker Hannifin (Australia) on twelve (12) 125mm crustbreak air cylinders of similar type to those used at Portland. The purpose of the trials was to determine which factors had the greatest effect on rod and piston seal life and to test alternative Parker Teflon seal units. The opportunity also existed to test other components such as piston rod material, rod/shaft connections, springs, stroke speed, cushioning and grease types. The accelerated trial was conducted at 40 times the normal plant frequency rate. Parker supplied facilities, designed equipment, and provided personnel to monitor performance during the time. The author specified operating conditions to which the cylinders were exposed, specified duplicate monitoring equipment, supervised independent validity checks, conducted gauge capability tests, selected parameter settings for each run, collated/analysed the data and wrote thefinal repor t (85). 7.3.2 Procedure Six cylinders at a time were mounted in the basement of Parker Hannifin's workshop in Sydney (Figure 7-1). Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permissionfrom Portland Aluniinium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 257 FIGURE 7 -1 CYLINDER TRIAL EQUIPMENT TEST RIGS Figure 7-2 shows the features of a cylinder trial rig. Each cylinder was heated by an electric trace in a frame which had its own individual temperature controller. The sensing points were thermocouples on the barrel and in a hole in the bottom block drilled immediately next to the rod seal unit so as to correctly monitor the actual temperature to which rod seals were exposed. The two temperatures followed similar trends. This duplication assisted in a few cases of thermocouple failure. Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permission from Portland Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 258 FIGURE 7 - 2 CYLINDER TRIAL RIG CYLINDER HEATING IACKET [THERMOCOUPLE CHAMBER FOR ALUMINA AERATION FLUIDISATION INSPECTION PORTS AERATION CHAMBER MECHANICAL STROKE COUNTER Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permission from Portland Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 259 The rigs were assembled and monitored 5 days a week by Greg Mansfield of Parker (under the direction of Gary Nauer). Special frames were constructed with the following features: (i) alumina aerated around the seal/mass, (ii) identical weight to that suspended under operating cylinders in the pots, (iii) eccentric masses to simulate misalignment seen on the feeders, (iv) temperature cycling to simulate hot/cold operation in the pots and anode effects, (v) on-line accelerometer on each cylinder to see the effect of cushioning over time, and (vi) pistons stationary for a short time prior to change of direction. Parameters tested included: (i) 2 different piston rod materials (10L45 and 4140), (ii) 2 different piston rod/shaft connections (M24x2 and stud), (iii) 2 different top cushion lengths, (iv) 2 different springs (stainless and Kempe high temperature), (v) 2 different piston seals (Viton and Teflon), (vi) 2 different rod seals (Viton and Teflon), (vii) 2 different stroke speeds (0.7s down/1. Is up and 2.0s/2.2s), (viii) 3 different suspended mass eccentricities (0mm, 5mm and 10mm off-centre) and (ix) 4 different greases (Tranz 414, A2M, B2 and B1). Although the list of variables is long, the prime test was for rod and piston seals. Seal life was expected to be affected by stroke time, eccentricity, grease and temperature but not by piston rod material, the rod/shaft connection, cushion length or spring type. Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permissionfrom Portlan d Aluniimum Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 260 The equipment comprised a large compressor, on-line air dryer (for -40°C dew point), 16 pen chart recorder (for temperatures), alumina circulation system, seal leak detection equipment, CRO (for accelerometer monitoring of cushioning) and high purity air filters. About $80,000 worth of equipment was used in the trial. A seal failure was deemed to have occurred if the leak rate was greater than 6 1/min for three consecutive days. The reason for the "consecutive days" was that the leaks tended to change as the cylinder cycled, possibly due to side movement or uneven seating of seals from time to time. (This is not unusual in the plant either, as operators see ore leaks come and go on a feeder.) Of the 12 runs, 10 failed due to rod seals and 2 due to piston seals. To ensure that there was complete candour, Russell Overhaul (with assistance from Dr Hugh Stark) from Unisearch (University of New South Wales) was engaged to randomly monitor the equipment unannounced. In addition, Russell Overhaul independently checked instruments with high precision calibrated instruments to ensure that results were valid. All Unisearch information cross-checked correctly with Parker monitoring throughout the trial. After each run, the cylinders were jointly stripped and inspected by Greg Mansfield and the author. The actual eccentricity of the mass was measured with a plumb bob, and all relevant dimensions of seals and components were measured in a detailed post mortem. The trial took 6 months to conduct, and analysis of data took about 200 man hours. Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permissionfrom Portland Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 261 7.3.3 Results Table 7-1 and Figure 7-3 show the results of the trials. Run 5 failed due to thermosetting of the "O" ring of the rod seal housing, and Run 6 from incorrect installation of the seal unit. 7.3.3:1 Sensitivity Analysis Comparisons are made where most factors are held constant except one parameter viz. stroke time, seal material, eccentricity. (i) Fast Speed - Viton The data for Runs 11 A, 11B, 22A, 22B and 6 are shown in Figure 7-3. It was concluded that the longer the time that the seals were exposed to high temperature, the shorter their life. The reason for the large differences in temperature exposure was due to random periods that the cylinders were exposed to simulated anode effects or simulated hot pots during the trial. (ii) Eccentricity Runs 11A and 11B had no eccentricity and 22A and 22B had eccentricity of 10mm; equivalent to Portland plant conditions. At higher temperature exposures (40 hours at 180°C) the eccentricity had no effect, but at low exposure times (less than 10 hours) there was a large increase in life for less eccentricity viz. 50%. (iii) Temperature versus eccentricity Temperature exposure had the greatest effect on seal life viz. changing exposure from zero hours to 55 hours changed life by the equivalent of 4.5 years. At low temperature exposure, eccentricity was the next most important factor viz. changing eccentricity from 10mm to zero increased life 1.5 years at low temperature (Runs 1 IA and 22A). Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permissionfrom Portland Aluminium Technical Manager. 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BES T VITON , / CO CJ CJ UJ H CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 264 (iv) Slow Speed - Viton All the slow speed runs had temperature exposure in the region 40-50 hours over 180°C. As eccentricity reduced, life increased viz. Runs 5, 3 and 4 had 10, 5 and 3 mm eccentricity respectively and resulted in 0.8, 2.8 and 3.0 years life. (v) Slow V's fast speed - Viton For similar high temperature exposure, comparison of Runs 11B/22B and 5/3/4 showed slow speed gave longer life; about 2 years at 5mm (or less) eccentricity. (vi) Fast speed - Teflon The lower eccentricity of Run 12 gave longer life (6 months) compared with Run 21 for similar temperature exposure. The wear rate on Run 12 was exceptionally good for the period it ran prior to thermosetting of the seals. The reason was unknown. (vii) Slow speed - Teflon The eccentricity for both Runs 13 and 23 was identical, but Run 23 had much higher temperature exposure time (40%) 12% more rod seal wear (as measured by seal thickness) probably due to the greater temperature exposure. This illustrates that greater temperature exposure weakens the seals. (viii) Slow V's fast speed - Teflon It is unclear of trends for Teflon with respect to fast and slow speeds from the test data due to the compounding effect of significant differences in temperature exposure across Runs 12/21 versus 13/23 respectively. (ix) Viton Versus Teflon Teflon seals were superior to Viton. Extrapolation of the Viton line for Runs 11 A, 22A, 11B, and 22B suggested a 2 year longer life for Teflon when over 50 hours above 180°C. Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permissionfrom Portlan d Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 265 7.3.3:2 Teflon Versus Viton Seals Initial examination of total life indicates that the new Teflon rod and piston seals proved unsuccessful relative to existing designs. The teflon seals had a silicon "O" ring to press the seal against the piston rod or barrel. The silicon "O" ring hardened and formed flat faces so that there was no pressure on the seal. This provided a route for air to escape. At the time of failure of the energiser, the seals were expected to have significant operating life still available. Note that silicon has a lower recommended operating temperature than Teflon viz. 260°C versus 300°C peak (87,88,89). One must consider "O" rings as well as the seal material when designing a cylinder. Stainless energisers are a safer design for high temperature applications (Section 8.4.6). From Figure 7-3, one can see that the Teflon seals were exposed to generally a much longer period of high temperature. This will have affected their life. Prediction of life at lower temperature was not possible due to the restricted number of runs with the budget and time available. Despite this, one can assume that the Teflon would have been superior to Viton if one compares the life of Viton at similar periods at high temperature. Note that independent feeder plants use Viton, so there is a chance that cylinders may not give good life in some potlines that have high background temperatures e.g. Portland or T8. 7.3.3:3 Temperature There is a clear result that exposure to high temperatures (for similar other conditions) drops seal life. These results were not unexpected. In the plant, several feeders from the same pot have been seen to fail within shifts/days after a pot becomes very hot (over 970°C). In order to optimise seal life, the periods of high temperature exposure need to be minimised and the design of seal/bush/scraper need to be addressed. Refer to Section 8.4.2 for discussion on methods to reduce the effect of high temperature exposure on seal life. Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permissionfrom Portlan d Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 266 7.3.3:4 Speed One would have expected failure of Runs 3 and 4 after about 3 months based on the results at the fast speed at similar temperature exposure. However, they failed at over 2 years. An 8 factor improvement was seen for this condition. This suggests a cubic relationship between speed and wear rate of seals which agrees with theory of frictional wear. 7.3.3:5 Eccentricity The effect of designing out a 10mm eccentricity relates to about a 50% (or 18 months) improvement in rod seal life at low heat exposure and fast speed, but no improvement at high temperature exposure. At slow speed, the trends for eccentricity were reversed at high exposure. However there was only minor difference in eccentricity for this comparison (viz. 3 and 5mm for Runs 3 and 4), so this may be due to sample variability or other factors. The ideal way to test this factor is to repeat the trials at 2 exposures and 2 eccentricities. This was not possible due to lack of time and money. Run 4 failed earlier than Run 3 due to thermosetting of the 'O' ring i.e. the energisers behind the seal hardened and lost their spring effect to push the seal against the rod or barrel. As there was effectively 40% of the seal still available, it is expected the rod seal should have lasted at least 40% longer (or 3.2 years). Thus it would be better than run 3 at slightly worse eccentricity (Figure 7-2 has Run 4 located at its assumed life). Hence, it was concluded that there is better life for less eccentricity with the difference being more noticeable at lower temperatures. Benefits of up to 2 years are possible. Improving alignment by increasing bushing length at the rod seal, maximising bush length of the feeder/crustbreaker and using spigotted joints are all useful actions to minimise the effect of this factor (Section 8.4.4). Independent feeders have less eccentricity than integrated feeders, so will have longer seal life (Table 8-III). Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permissionfrom Portlan d Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 267 7.3.3:6 Grease Tranz 414 was used in all runs of Viton except for 2 runs with A2M. B2 was the only non-insulating grease. Seven of the 12 runs showed grease breakdown. There was no pattern between temperature profile and lubricity at failure. However, A2M did not suffer from breakdown despite a long exposure to high temperatures. Technical data from the supplier suggests this grease should be superior to the others due to 60°C higher temperature rating and smaller particle size. This grease was later tested to be better than Tranz 414 in the lubrication investigation discussed in Section 7.4. 7.3.3:7 Piston rod/shaft connection There was no evidence of chrome lift-off on any piston rod, and no fatigue failure of connections. This was probably due to generally lower than normal acceleration rates on these cylinders (compared to plant experience) and low cycles encountered - all except one run were under 3 years of service. 7.3.3:8 Top cushion length The effect of top cushion length was not conclusive due to the small sample size and no piston rod failures to illustrate the differences between the two designs. 7.3.3:9 Spring design One stainless spring fractured a couple of turns from the top (Run 11 A). This type of failure has not been seen before, so it was assumed it may be related to having the spring compressed length too small in the retracted position on this cylinder. One cylinder (Run 4) started to shrink which is common in service at Portland. There were no faults with the Kempe high temperature springs. It may be concluded that the Kempe springs were superior to stainless. Confidential: Written permission required by both parties to publish or reproduce outside j\lcoa and Parker Hannifin. Alcoa permission from Portland j^luminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 268 7.3.3:10 Deceleration rate Deceleration rates were monitored daily on each test viz. 5 days per week plus spot checks by Unisearch. There was a trend that deceleration rates tended to cycle up and down for several days at a time. There was no trend that they increased over time. Deceleration rates increased several fold when the cylinders were exposed to high temperatures. This is probably due to the very large drop in air density with increasing temperature. For example, at a pressure of l,000kpa, a change in temperature from 100°C to 200°C drops air density by a factor of 40^. This drop reduces dramatically the cushioning as there is no "body" in the air to stop the piston and metal-to-metal contact occurs. This will have the biggest effect when the cylinder is retracting as the dosing cup is hit with no cushioning on an AEDD feeder (Figure 1-9). This may explain an observation by operators that feeders get "stuck" on hot or new pots. What they often hear is a large "bang" and assume it was due to the plunger getting stuck in the crust. It is suggested this noise is mainly due to the higher deceleration force at the higher temperature. Once again, the effect of pot conditions on feeder operation is evidenced. There appears to be little that can be done to avoid this happening other than the obvious action of reducing the temperature of the pot. 7.3.3:11 Air leak rates A small leak varied from cycle to cycle. This is why a run was only terminated if the leak was excessive for readings on three consecutive days during the trial. If plant air checks are conducted on a plant to determine if air leaks are excessive, one must be careful that decisions are not affected by this variability. It is preferable to conduct a gauge capability study to determine at what levels of air leak that feeder changeout is required. A leak rate of over 301/min per feeder was found to be critical after these tests were completed at Portland. Air leak rates at room temperature were up to 7 times higher than at operating temperature, but this ratio varied considerably. Portland is now monitoring air leak tests Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permissionfrom Portland Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 269 into and out of the workshop and overhauling only those that have leaks over 30 1/min. Previous to this cylinder trial, if there was any leak at all, the cylinder was completely overhauled. Due to the effect of temperature, this old procedure led to overservicing of cylinders and excessive repair costs. The old procedure is common at most plants. (Refer to Section 8.4.4 for discussion on non-routine cylinder overhaul cost savings). The 30 1/min plant acceptance limit at Portland is much higher than the acceptance criteria used in the trial viz. 6 1/min. The plant acceptance condition is based on economic grounds, but the trial value was based on Parker's standards of a satisfactory seal. In plant conditions, it is possible to accept higher leaks if the leak does not cost too much air/money or cause other process problems such as ore leaks. The results of the trial on Run 22A (Figure 7-4) suggests that leak rate may not decay very fast and that the leak rate is not a linear relationship to seal life. Hence, the best changeout procedure appears to be to replace only on a needs basis rather than on a time basis (Section 9.4.3). FIGURE 7 - 4 CYLINDER 22A ROD SEAL LEAK RATE LEAK RATE (1/min) 80 80 70 70 60 60 50 - h- 50 40 40 30 30 n OO 20 20 D -J ULJ a D - • 10 — - 10 - nn c - 0 nHffl UUUn • I 1 u J 0 0.5 1 1.5 2 2.5 3.5 EQUIVALENT YEARS IN SERVICE Note: 200,000 cycles is equivalent to one year service in Portland pots Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permissionfrom Portland Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 270 7.3.3:12 Insulation Insulation dropped to zero over the life of every cylinder trailed. This has also been confirmed in the plant (Section 8.7.2). Checks on all types of greases used in the trial (except B2) showed that they were insulating initially but all broke down during the trials. It is assumed that the small amount of rust and material that gets through the filters in the airlines gradually causes cylinders to short out electrically. Note that a very fine filter was used in these trials (10 micron versus 50 micron at Portland) to minimise any chance of this happening, yet the insulation still dropped to zero. Section 8.7 discusses insulation further. 7.3.4 Cylinder Trial Summary Table 7-II summarises the data as discussed above. This shows that the importance of parameters were temperature, speed and eccentricity (in the order most important to least important). Table 7-in lists recommendations for design and use of cylinders. Portland and Parker have actioned many of the recommendations. Almost all trial results have been substantiated in plant conditions at Portland. These plant scale results are discussed further in Sections 8.4 and 8.5. This verifies that the trial conditions were comparable to plant conditions and is testimony to the effort put into planning and implementing the trial. Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. j\lcoa permissionfrom Portland Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 271 TABLE 7- II PORTLAND/PJ\RKER CYLINDER TRIAL DATA SUMMARY ROD SEAL HIGH TEMPERATURE CYCLES ECCENTRICITY SPEED LIFE MATERIAL EXPOSURE (X1000) Viton 5 hrs Zero Fast 5 years 1,000 Normal Fast 3 years 600 40-50 hrs Zero Fast 1 year 200 Normal Fast 1 year 200 40-50 hrs Zero Slow 3 years 600 Normal Slow 1 year 200 Teflon 50-60 hrs Zero Fast 2 years + 400 + Normal Fast 2 years + 400 + 50-60 hrs Zero Slow 2 years + 400 + Normal Slow 2 years + 400 + IMPORTANCE PRIORITY PARAMETER CHANGE DROP IN LIFE Temperature 50 hours exposure over 180°C 4 years Speed 50 hours exposure over 180°C 3 years Eccentricity zero to normal 1.5 years Teflon was better than Viton at high temperatur e Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. .Alcoa permissionfrom Portlan d Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 272 TABLE 7-ID RECOMMENDATIONS FROM CYLINDER TRIAL * Temperature, speed, eccentricity in priority of importance. CYLINDER SUPPLIER * Reverse bush and rod seal to have seals near air cavity. * Soften viton seals to accept some thermoset. * A2M is best grease. * Temperature stability of energisers is important as these were the first cause of failure for 6 of 12 runs. * Rod seal/bush installation can be installed incorrectly easily. - need keying to prevent errors at installation. * Teflon seals better than viton at high temperature. CYLINDER USER * Temperature - use old feeders for new pots. - vent air leaks and delay feeder changeout on hot pots. * Speed - slow the feeder down as much as possible. * Eccentricity - install better guides and bushes or universal joints. * Use A2M grease not tranz 414. * Use more grease on overhaul. * Check air leak rate in workshop on overhaul and accept a small air leak. * High temperatures cause high G forces - plungers do not get "stuck". Cushion settings and speeds do not change with time, so workshop settings can be used for records and decision making. * Insulation not necessary for new cylinders. Confidential: Written permission required by both parties to publish or reproduce outside Alcoa and Parker Hannifin. Alcoa permission from Portland Aluminium Technical Manager. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 273 7.5 LUBRICATION INVESTIGATION 7.5.1 Introduction A key factor in the life of piston and rod seals is the type, quality and nature of lubrication. In many plants no lubrication is used at all - even during overhaul of the cylinder e.g. most Alcoa smelters. Some plants have continuous oil dispersant systems in the air line e.g. T17. Most smelters use a lubricant only during overhaul e.g. plants using Atlas Copco or CPOAC cylinders (Table l-II). At Portland, analysis of rod seal life of identical units in Terry and Parker cylinders over several years shows a mean life of 550 months and 1080 months respectively using Weibull analyses. One possible cause of this difference is grease. Terry cylinders have used Magnalube and Parker have used Tranz 414 until this grease investigation. Workshop experience is that Magnalube breaks down more than Tranz 414. If the lubricant and/or the wiping system are not adequate, the lubricant can break down and form a grinding paste that can seriously affect seal life. Lubricant breakdown products can also be vented from the cylinder and block the exhaust muffler. This in turn causes back pressure on the cylinder which affects the net pressure drop across the piston. For integrated feeders this results in a speed change, change in shot size variability and reduced pressure at the crust. Blocked feeder holes can develop (refer Sections 3.5.3 and 4.5.3). Discussions with several seal and cylinder manufacturers showed varying attitudes to lubricants. Almost all suggested that use of a good grease could enhance significantly the life of seals, but, except for Atlas Copco, none had trialed alternatives or had advice on the best type to use. Based on this lack of information and difficulties in finding any independent expertise, CETEC Chemical Technologies (Dr. Vyt Garnys) was engaged to analyse a range of greases and to co-ordinate analyses by Shell and CETEC under the author's direction. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 274 Technical and laboratory assistance was given by John Quinlivan, Rainer Schade and Mauro Del Frate from Shell. An investigation of alternative greases (plus no grease) was conducted by the author using a Kepner Tregoe Decision Analysis. This method lists the "musts" and "wants" of good lubricants and rates the alternatives based on weighting of importance of each "want" and the relative merit of each alternative. Eleven (11) alternative lubricants were investigated after initial screening of products offered by Shell, Mobil, Optimol, Tranz Performance Lubricants, Skega and Magnalube. Except for Mobil, at least one grease was used from each company plus use of no grease at all (which is the situation in several Alcoa plants). No suitable Mobil lubricants were available in Australia so they could not be included in the trials. Except for Skega GPSO (which is a suspension) all alternative lubricants in the trial were greases. Detailed results are available in Kissane (85), but for simplicity only the relative ratings used in the Kepner Tregoe Decision Analysis are listed in the summary of results shown in Table 7-IV. Ill to > > > o ,_ o ,_ ^_ (B S n a o < I TJ TJ "g TJ o o o g UJ c c c ^ OC o UJ ra CO > > > eo •» ° TJ 3 £ 1 o o E - CM IO a: O TJ .C TJ IA O co UJ c 3 O < c is 3 £ s > >- > TJ a CO IO in TOJ « »F O 3 TOJ c c ra ra >- > >- ra1 £ TJ CO CN CN CN CD !A ToJ T8J £ O c c co 4= ro to .-£ ro (-, ,—. > > >- s « o ° C TJ Ur TJ °C4 ° > >- > 2 in CD 'E I -Si 1 co g o 8° o o S o CN *^ >• >• > TJ TJ £ TJ ^ Ol 8^ o o 3 o CN ^ 5 co JJs ro 0 T- T- > > >- TJ TJ !c TJ 2 O) - o o s o EEC >- >- > 1 lljflo $ "3- E TJ TJ r tl ° <- r- hi • 8 S * 8 UJ S3 3 o V >- > PJ CT> ° E TfOl TtSO ° r- _i O O S O CO < (fl c c c Z o f H -* 1 > >- > 'E ra co o ro ,—, CO 00 O TJ TJ j TJ o N° O O J2 O EC E < IN >- > > 00 0> CNOIO) c TJ TJ .E TJ ^£ CO < s 9 2 5 o a> oo ^C CD CO CN T- io m CO e a. S' 8, » E o O I s O) CO A „ i? £ S | o I 8| T J I &»• SJ^ ° % c Ol v. 11! iff t 3 o o o e w • 3 « ° i— i— Due to the complex nature of the duty of a pneumatic cylinder, a range of tests were used to simulate long term medium temperatures, short term peak temperatures, oxidation, side load, and grinding nature of grease residues. It is useful to briefly explain the methods of analysis and the equipment used. (i) Differential Scanning Calorimetry (DSC) A 5mg sample was heated in an aluminium pan at the rate of 10°C/min in air at 70kPa to 400 C. The heat flow into and out of the sample (using an empty pan as a reference) was measured as a function of temperature. Dupont DSC analytical equipment was used. (ii) Thermogravimetric Analysis (TGA) A lOmg sample was heated in a platinum boat at the rate of 20°C/min in air to 1,000°C. The amount and rate of change in mass of sample was measured as a function of temperature. Dupont TGA analytical equipment was used. (iii) Penetration (ASTM D217) The consistency of grease was measured by the penetration of a standard cone after working the grease 60 strokes. The penetration was determined at 25°C by releasing the cone assembly from the penetrometer and allowing the cone to drop freely into the grease for 5s. A coarse penetrometer was used. Note that greases are classified by a "NLGI (grease)" consistency number according to the penetration measured. (iv) Dropping Point (IP396/93) The dropping point of a grease is the temperature at which the first drop of material falls from a cup. It is the temperature at which the grease passes from a semisolid to a liquid state under the conditions of the test. Samples were heated at both the rate of 0.2°C and 1°C per minute. A Mettler Drop Point Tester was used. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 277 (v) Sen wing Reibung Verschleiss device (SRV) (ISO 9001/9003, DIN draft 51/834) The Schwing Reibung Verschleiss (Vibration Friction Wear) device uses a ball and plate, or roller and plate mechanism. The test specimens were mechanically loaded at a given frequency, load and stroke. Tests was conducted at set temperatures up to 180°C. Friction force was measured continuously and friction coefficient was calculated and plotted on a chart recorder. This device was used with Teflon as the plate material. The tests using a ball simulated "extreme pressure" (which is applicable to wear of a bush) and tests using a roller simulated seal wear. This device was considered to be the best instrument that exists to measure seal friction coefficient under a range of operating conditions. It was used to compare the wear scar of different types of Teflon (Section 8.4.5). 7.5.3 Kepner Tregoe Decision Analysis Procedure The Kepner Tregoe decision analysis considers MUSTS and WANTS (Table 7-IV). MUSTS are "not negotiable" requirements; if the alternatives do not meet these conditions they were scrapped. In this study, all 12 alternatives passed the MUSTS. WANTS are desirable outcomes but of different levels of importance. The WANTS were listed and weighted by importance from 10 (best) to 1 (worst). Each alternative was tested against the WANTS. "Weights" were placed against each WANT to compare the importance of each WANT relative to the other WANTS. The "weightings" were listed prior to any comparison of alternative to avoid any influence of "favoured" alternatives. Each WANT was then examined to establish the best parameter to quantify the 'WANT*. For each alternative, the actual analytical result is listed and a "rating" across alternatives established their relative benefits from 10 (best) to 1 (worst). Some 'WANTS' had no data because there was (a) no data available, (b) no reasonable method to analyse for it, or (c) no equipment was available to conduct such a test. It was considered at this stage (and later assessment at the end of the decision analysis) that the CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 278 results of the WANTS which had no data would not have changed the outcome due to their low "weighting". Once WANT "weightings" and "ratings" were established, these were multiplied together and added to give a total for each alternative. This reduced the list from 12 to 4 alternatives (Table 7-V) For the final four, "adverse consequences" were examined for each to reassess if there was a critical element that was very weak thus making it sensitive to a particular extremes of operation. TABLE 7 - V FINAL LUBRICANT RATINGS LUBRICANT POINTS ADVERSE CONSEQUENCES RISK 1. Darina 359 unstable at high temperatures HIGH 2. TA2R2 357 untried, but high rating in all areas LOW 3. Titan A2M 345 untried except PA/Parker trials LOW 4. Skega 338 good record at Atlas Copco plants LOW 5. Aeroshell 323 6. TA2R1 322 7. Magnalube 310 8. Tranz 414 296 9. No grease 267 10. Stamina 262 11. Paste MP 260 12. Innertox 229 Confirmation tests for top four: SRV roller and Teflon plate, 1 hour at 180°C after 8 hours at 300°C. Final rating: 1. TA2R2 2. Titan A2M 3. Skega GPSO Darina... major failure - deleted CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 279 Of the short list alternatives, the first alternative (Darina) was particularly sensitive to spikes of high temperature so was a "high risk/high probability" combination. The other three alternatives (TA2R2, Titan A2M, and Skega GPSO) had no serious adverse consequences. Further comparative testing was conducted of heating each of the top four to 300°C for 16 hours then doing a 180°C roller and plate (Teflon plate) test on the residue. It was concluded that none dropped out from the list except Darina. This confirmed the suspected sensitivity of Darina and it was decided to reject it. This left TA2R2, Titan A2M and Skega GPSO for future trials on plant scale routine cylinder overhaul. Tranz 414 would act as a "control" grease, as this was the best of the greases already in use. 7.5.4 Discussion 7.5.4:1 Benchmarks Tranz 414 and Magnalube were used as benchmarks throughout the investigation as these have historically been used at Portland. Historically, Tranz 414 has been superior to Magnalube at Portland. It would be expected that Tranz 414 should have a higher rating to Magnalube in the laboratory tests and that other greases should be superior than both Tranz 414 and Magnalube. At the end of the decision analysis, Magnalube rated better than Tranz 414, but the final scores were not significantly different. Both were well behind the best lubricants (despite being considerably cheaper). Skega GPSO has been used by Atlas Copco in all their crustbreaking cylinders for many years. Atlas Copco have done considerable large scale testing of many greases at elevated temperatures. Skega GPSO was one of the three top options from the decision analysis. It showed good all-round ratings with few low-rating qualities. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 280 Thus, lubricants found to be successful in operating plants also fared well in the laboratory tests and decision analysis. It was therefore concluded from the benchmark comparisons of Tranz 414, Magnalube and Skega GPSO, that the laboratory tests and decision analysis should indicate a satisfactory rating of performance that would be expected to give comparable relative results in operating plants. 7.5.4:2 No Lubrication The advice received from Atlas Copco, Parker, Terry, CETEC and Shell was that grease must be used with Viton. Teflon has some lubrication ability, but Viton has none. This has been verified at Portland T9, T16 and T17. Two cylinder types have been used at T9 and T15, and neither had any lubrication on either Teflon or Viton. There was a major difference in life despite having identical assemblies connected to the cylinders. The option of "no lubrication" had interesting results. This had a better friction coefficient at elevated temperatures but was far worse than the others at temperatures under 180°C; almost three times worse than Tranz 414. It appears that "no lubrication" is not as good for normal operation, but is not affected by the residue grinding action when the seals are "cooked". It is believed that insufficient lubrication has been a major problem at Portland, especially from October 1991 to February 1993 when the amount was reduced still further because it was considered that it was a waste of money to put too much grease on. Seal life increased from the time that the amount of lubricant was increased (irrespective of the grease type). Figure 7-5 illustrates the effect of the amount of grease. Initially, grease was wiped on with a small spatula on the barrel, piston seals and rod seals. Then, in October 1991, this was reduced to using a small paint brush. In February 1993, the amount was increased to the maximum that would physically fit into any part of the rod seals/bush area and onto the piston seal/bush area plus the barrel was liberally covered. The rod seal life dropped by a minor amount on the first reduction, but increased by 45% when the amount was a maximum. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 281 FIGURE 7 - 5 EFFECT OF AMOUNT OF GREASE ON ROD SEAL LIFE (a) NORMAL AMOUNT (PRIOR TO OCT,91) AVERAGE LIFE: 450 days 100 -y. y' j- 80 50 20 10 8 = 1 ., 9 y* FaTk J re 3 = - Ceni (or 3d 2E M. 1B0. XAL MO. 1000L XKICl WOO. AVERAGE LIFE: (b) REDUCED AMOUNT (OCT,91^JAN93) 440 days 100 / s 80 ,ry y~ - 50 * + / r 20 .^ 10 >" r _« ,<"^ 0 = 1 -{ ^ TJ — * xJJ> " -'FoTU i = f 53 2 Cam por ad = 3 1 -k is. aa i DO aoo. 600D. 1 AVERAGE LIFE: (c) INCREASED AMOUNT (MAR93-JUN93) 660 days 100 80 50 _*:^ - 20 • \/ ' 10 f 5 ff = 1 .1> A •n = 73 4 '' y * • FoTh ire 1 a 1 3 / Certi tor 3d ; 1 E l> . ' io. aa. ai. 100. an. soa. tace. Total Opwotfng Tim* (©aire) Notes: (i) Basis: Magnalube grease in Terry cylinders using FEC part #423. (ii) Vertical axis is "cumulative percent occurred''. (iii) "Censored" refers to replacements that were not due to failures e.g. preventative maintenance. CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 282 Note that this comparison was for the same seal and grease, and there had been no changes to bath penetration. Stroke speeds dropped between (a) and (b), so this may explain why there was not an even greater drop in life when the amount of grease was initially dropped. The slower speed and friction drop compensated the potentially large drop in life from less grease. Conversely, the expected increase in life of seals from the lower speeds did not eventuate probably due to the reduction in the amount of grease at the same time. This example illustrates the compounding effect of making two changes at the same time. It could be suggested (after the fact) that doing two things at one time was unwise, but one must recall that the decision to reduce the amount of grease was considered by Parker and Terry to have insignificant effect on seal life at the time. Only after the cylinder trials and initiation of investigation into lubricants was the possibility that this may be significant was contemplated. Thus, controlled experiments can be affected by one's perception of what is significant. 7.5.4:3 Lubricant Options If a lubricant is to be used, it is important to select one that not only operates well at low temperature, but does not form a grinding paste at high temperatures. This requires it to withstand oxidation, "bleed" (flow away from its original location by gravity) and high pressures. When these considerations were examined, the lubricants tested performed quite differently. Some were better at high temperature, others take a long time before breakdown, then suddenly totally disintegrated. If insufficient lubricant is used, the higher friction forces cause rapid breakdown (of even good greases) due to shear. It is necessary to have a reservoir always available so the lubricant in the wear area can be replenished easily. For a vertical cylinder this requires a lubricant that is thin enough to replenish the work area yet not too thin to fall to the bottom of the cylinder over 5-7 year period even under extreme temperatures. To select a lubricant that handles all the desirable requirements is quite a challenge. In Shell's opinion, selecting a lubricant for smelting pot alumina feeders is one of the most CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 283 difficult duties they have ever undertaken due to these complex issues. The solution to the problem is state-of-the-art lubrication. However, not to use grease will give poor seal life (especially for Viton). Shell advised that copper and iron are catalysts to the breakdown of grease. Hence use of bronze wipers (as used by Portland and T8) will affect grease life. Teflon wipers are used on Atlas Copco, CPOAC, Lindberg, and Scheffer cylinders. Portland changed to Teflon late in 1993. 7.5.5 Conclusions and Plant Scale Results The recommended lubricants were 1.TA2R2 2. Titan A2M 3. Skega GPSO The top two are made by Tranz Performance Lubricants (Sydney), and GPSO is marketed by Atlas Copco (Sweden). These were trailed in the field since July 1993 using liberal quantities for each cylinder overhaul. Compared to Tranz 414, cylinders using the above three lubricants have doubled their rod seal life. Thus, the conclusions from this study have proved to be correct in operating pots. As a result of this lubricant study, Tranz Performance have initiated enhancements on TA2R2 to develop TA2R7. Tranz TA2R7 has been used at Portland since early 1994 with very good results. It is considered to be the best lubricant used to date, but it is fairly expensive at about $10/overhaul or $300/kg. However, the cost is justifiable if it increases cylinder life even by such a small amount as the average cost of a cylinder overhaul is about $400-500. More time is required to do a cosl/benefit analysis comparison between TA2R7 and TA2R2 (which is considered to be the best of the other three). CHAPTER 7 PNEUMATIC CYLINDER INVESTIGATION PAGE 284 7.11 MAJOR FINDINGS OF PNEUMATIC CYLINDER INVESTIGATION (i) Main factors affecting rod and piston seals (in order of priority from most important) are (a) temperature, (b) stroke speed, and (c) eccentricity. (ii) Teflon is superior to Viton with respect to wear and temperature sensitivity. (iii) Good lubricants for pot feeder pneumatic cylinders are TA2R2, Titan and Skega GPSO. (iv) Viton seals need lubrication. (v) Lubrication can drop friction coefficient by up to three times that of no lubrication (when under 180°C). (vi) Grease rapidly breaks down if insufficient quantity is used. (vii) Copper and iron catalytically break down grease. 7.12 RECOMMENDATIONS FROM PNEUMATIC CYLINDER INVESTIGATION (i) Use adequate quantities of high temperature lubricants for feeder/crustbreaker cylinders. (ii) Use Teflon seals. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 285 CHAPTER 8 EXTENDING FEEDER UNIT LIFE 8.1 SUMMARY From the frequency of each type of failure of feeder/crustbreaker units, an analysis was carried out of ways of extending their life. Ore leaks on integrated feeders can be addressed by a minor alteration to a feeder to allow leaking rod seal air to escape from the pot so as not to affect the pot operation. A system to measure air leaks in the potrooms and in the workshop monitors excess air usage and poor feeder quality so that these do not get out of control. Opportunities exist for plants to extend life and reduce costs by accepting limited air leaks at overhaul, not overhauling feeders from off-line pots and not requiring insulation for newly purchased cylinders. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 286 8.2 INTRODUCTION Following research into the dosing system (Chapter 2), plunger wear (Chapter 5) and cylinders (Chapter 7), a clearer understanding of the components of feeding and crustbreaking units has been established. This Chapter will consider in detail the causes of failure of cylinders and feeding assemblies that lead to the feeder being removed and how to minimise the frequency of these failures. Firstly consider the interaction between operators in the potline and the repair tradespersons. It is useful at this point to highlight an important difference between "reason for removal" and "mechanical fault". "Reason for removal" is why the observer believed the feeder was not performing (such as not feeding or feeding too much). The observer cannot see the internals of the feeder so cannot state the reason was due to, say, a spring failure or to the spool jamming. There may be several mechanical faults that can contribute to a feeder being removed. For example, alumina leaking unmetered into the pot can be caused by either cylinder failure or by dosing unit failure. It is important that one is aware of the nomenclature on removal reasons as one may spend a lot of time working on the wrong things. As little testing is conducted on shot size accuracy of overhauled or operating feeders for both integrated and independent designs, the usual reasons for removal are: (i) a significant overfeeding or underfeeding situation which is visually obvious to an observer, or (ii) if the pot is not operating properly and the feeder is diagnosed to be potentially at fault. This diagnosis is often fairly subjective so feeders are frequently removed inappropriately. Good troubleshooting procedures should be used before feeder removal as feeder repairs are expensive and expose people torisks of falls and exposure to heat, dust and fumes. Removal should only be used when one is sure the feeder is the cause. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 287 Section 8.3 links the reason for removal and the mechanical faults. Subsequently, Sections 8.4 to 8.8 discuss the mechanical faults in more detail. 8.3 CAUSES OF FEEDER FAILURES are. The causes of feeder removal of 11 integrated feeder plants in 1991 represented in Table 1-V. These are due to ore (alumina) leaks (44%), pointed or short plungers (20%), arced (13%), pot offline (13%), broken piston rods and dosing spools (8%) and not feeding (2%). In recent years, removal of feeders because of excessive cylinder bypass (piston seal leaks) has been added to this list. The reasons and extent of failures is quite different in each plant despite having very similar designs. Clearly, there are different problems at each plant, however, this presents an opportunity to optimise the designs, maintenance and operations of feeders by exploring the strengths and weaknesses between different smelters. Independent feeders do not generally suffer from ore leaks but the other causes do occur (though less freo^erAV, than for integrated feeders). The major causes of failure of are independent feeders leaking piston seals (cylinder bypass) and worn plungers. Rod seal leaks are important for air usage reduction, but do not pose a problem to feeder operation as the leaks are usually vented well and do not have many downstream effects (such as ore leaks). Rod seal leaks mainly create a cost penalty in independent feeder plants rather than a cause of feeder removal. As there are no springs, the dosing unit failures for independent feeders are associated with jamming of the spool and failure of the spool shaft (or bush that supports it). The author has limited data on spool failures in independent feeders so there is minor discussion on this subject. However, the discussion on spool dosing accuracy (already discussed in Section 2.5.7 and 2.5.8) and integrated feeder spool failures (Section 8.8) may be relevant to those interested in this subject. CHAPTERS EXTENDING FEEDER UNIT LIFE PAGE 288 Cylinder life and plunger wear remain the main causes of feeder unit failure in most plants - integrated or independent. Consider now the "reasons for removal" and trace them to their "mechanical faults". 8.3.1 Ore Leaks An ore leak is where alumina (ore) pours into the pot unmetered and potentially causes the pot to become mucky. This is usually seen as a dribble of alumina flowing from inside the feeder chute or, if very bad, at the kidney plate (Figure 1-11). It can be caused by several situations: (i) Rod seal air leaks The rod seals are located where the piston rod goes through the end of the cylinder (Figure 1-20). These leaks aerate the alumina in the dosing unit so the material does not seal where the spool seats on the fixed face of the assembly (Figure 1-10). If the air volume is too large, the dosing unit pressurises, causing alumina to come out of the kidney plate area like a curtain around the feeder chute. The reason the air goes downwards instead of upwards is that most integrated feeders have a solid joint at the top mounting flange where there is usually an insulating gasket or there is a metal-to-metal mating of flanges. This prevents air from escaping from the top of the cavity where the feeder is mounted. Most plants have a vent pipe from this cavity but this pipe can block if alumina is carried out with the vented air (which is usually the case as fluidised alumina has a 7° angle of repose and the level of alumina in the superstructure is above the level of the vent pipe). Section 8.4 discusses rod seal leaks in detail. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 289 (ii) Spool Jammed The overlap between the dosing cup and the spool closing face is only 2-5mm on an AEDD feeder (Figure 1-11) and is less than 2mm for independent feeders. Either partial or complete jamming can occur if the spool is not made correctly (refer Section 2.5.8 and Figure 2-16), if there is mechanical damage at the dosing cup, if mechanical failure causes the spool to be bent or if there are foreign objects fouling the sealing face. The common cause of damage is where feeder mounting bolts become loose and fall into the alumina and foreign objects subsequently try to pass through the tight clearances of the dosing unit. Depending on the alumina supply system, there are other potential causes of foreign objects getting into the alumina. Plants often have mesh installed upstream or just prior to the inlet to the feeder (T7), but these are commonly not inspected or cannot be inspected (for example, T7). Pechiney developed a sophisticated filtration system to prevent foreign objects getting into their hyper- dense phase alumina supply system as part of their AP30 pot design (51). Section 8.8 discusses dosing failures in more detail. (iii) Spring Failure Spring failures are not relevant for T8 or independent feeders. If a spring fails, it does not force the spool flat against the horizontal sealing plate (Figure 1-11). As the spring only extends for a feed every 3 minutes or so, it is compressed for about 98% of the time. This is a most difficult task for any spring let alone a 302SS or 304SS spring commonly used in Alcoa AEDD feeders viz. Portland, T2-T7 and T9-T17. The spring will creep over time, but is seriously affected by excursions in temperatures. Section 2.5.3 explains the effect of spring quality on dosing accuracy and Section 8.8.3 discusses what can be done about preventing spring failure. For independent feeders, ore leaks do not usually occur because: (i) there is no spring in the dosing unit to fail, CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 290 (ii) dags do not affect dosing cylinder stroke as the dosing unit is separate from the crustbreaker, and (iii) generally the cylinders vent into the pot cavity in large vent passages so there is little chance of back pressure. When one considers that in 1991 44% of integrated feeders were removed for ore leaks, this attraction of independent feeders is a very significant advantage and a clear contributor to better life for independent feeders. 8.3.2 Pointed and Short Plungers As seen in Table 1-V, after ore leaks, the next most frequent failure (20%) of integrated feeders failures is the combination of pointed and short plungers (which is due to excessive penetration in the bath). No data is available for independent feeders except D12 which had 85% of failures for this reason. Most independent feeders use stainless plungers to address this problem. Plunger wear is discussed in Chapter 5 where not only the material of construction is discussed but a range of other things that can be done to improve life viz. less penetration, less time in the bath, do not feed on anode effect. 8.3.3 Arcing The feeder and superstructure should be electrically insulated to prevent an electrical circuit being made between anode and cathode if the plunger penetrates the bath. Although for normal operation there is only about a 4V DC voltage drop across a pot, it is possible that this may reach 100V DC in anode effect. At the higher voltages of anode effect, arcing can occur between metal parts that are close together. This is usually where the dosing unit part of the assembly touches therim o f an AEDD feeder (Figure 1-11) or similar location on other feeder designs. The other locations for arcing are at the rod seal area of an air cylinder or, less frequently, at the piston of the cylinder. About 13% of feeders changed in the integrated feeder plants surveyed in 1991 were from arcing. It is rare that independent feeders suffer from arcing due to good mounting and lower end insulation. Arcing is discussed further in Section 8.7. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 291 8.3.4 Fatigue of Piston Rods and Spools Broken spools and broken piston rod comprised 8% of integrated feeder failures (Table 1-V). Both can be related to excessive piston speed, but both can be caused by inferior design. Side loading of the plunger when it is fully extended puts considerable strain on piston rods where they join the piston (Figure 1-20). Piston rods generally fatigue at this location. Note that only Portland and T6 of the 11 plants surveyed in 1991 had broken piston rods. Since this time, D12, T8, T15 and T17 have made enquires to the author about problems in their plants from broken piston rods. This failure mechanism may be greater than first indicated and is probably plant specific. Section 8.6 discusses broken piston rods. AEDD feeder spools can fatigue at the horizontal face below the spring (Figure 1-11) or they may wear at the top pipe section closest to the cylinder. Independent feeders can be affected by bent shafts or broken bushes that support the shaft above the spool, but generally do not fatigue due to slower speeds and lower driven mass than integrated feeders. Section 8.8.1 examines spool failures. 8.3.5 Cylinder Bypass Cylinder bypass is when the piston seals leak causing air to bypass through the cylinder (Figure 1-20). This a major cause of failure of independent feeders due mainly to proactive attention to reduce plant air usage rather than a factor that prevents the feeder from doing its job. For integrated feeders this historically has not been a major issue (as illustrated by the fact that in 1991 not one of the 11 plants surveyed listed it as a cause of removal. In the last few years following the authors feeder recommendations (Table 1-VH), Alcoa plants have looked closely at cylinder bypass and have detected very significant air usage concerns. The cause of cylinder bypass is generally piston seal failure (refer Section 8.5). The control strategy to minimise air usage is discussed in Section 8.4 as the choice of whether to change the feeder is often based on economic grounds provided the plant has sufficient air available. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 292 8.3.6 Pot Off Line This is the one cause of failure that most people want; 12% of feeders changes out in the surveyed plants in 1991. However, this research has found that these repairs are not cost efficient, they are unnecessary and they have a negative impact on pot operation, so there are considerable effort and savings by not doing feeder changeouts when pot superstructures are overhauled. This will be discussed in Section 8.5 as this is purely a cost based decision and not related to failure as such. 8.4 ROD SEALS 8.4.1 Causes of Rod Seal Failure Rod seal failure is one of the major causes of ore leaks in Alcoa pots and hence one of the major causes of feeder removal (Table 1-V). Examination of the cause of failure of 46 rod seals at Portland is seen in Table 8-1. TABLE 8 -1 CAUSES OF FAILURE OF ROD SEALS AT PORTLAND Worn 34% Cooked 24% Preventive maintenance 15% Alumina in cylinder 15% Broken piston rod 4% Other 8% 100% Sample size: 46 The priority order may be different at other plants, but the factors are often similar. Wear is due to temperature, speed and eccentricity as established in the cylinder trials (Section 7.3.4). The other causes relate to high temperature exposure (cooked), ingress of alumina past the wiper (alumina in cylinder), fatigue (broken piston rod) and other factors. Figure 8-1 illustrates schematics of some rod seal and bush combinations commonly installed. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 293 FIGURE 8 -1 CYLINDER ROD SEAL AND BUSH DESIGNS (a) CONVENTIONAL - SINGLE SEAL CYLINDER CAVITY WIPERl ROD SEAU SPACER BUSH 'O' RING (b) SPLIT BUSH ^^^1 1 Mr^. (c) REVERSED (d) TEFLON DOUBLE SEAL TEFLOl WIPER \EI\S, TWO SEAL CARTRIDGE CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 294 The causes of rod seal failure are discussed in turn below (with the exception of broken piston rod which is discussed in Section 8.6). This section only discusses rod seals, but one needs to also note comments in Section 8.5 on piston seals, as they are interrelated. 8.4.2 High Temperature Exposure Control The Portland/Parker cylinder trials (Section 7.3) confirmed that high temperature reduces seal life. There are several areas where operators or designers can contribute to cylinder life. 8.4.2:1 Pot and Feeder Design There is little that can be done about increasing the distance between feeder and the hole except on the design of new pots where it is best to use long stroke cylinders to take the plunger as far as possible from the crust. Integrated feeders have a crustbreaker stroke distance of 356mm (except for T8 at 470mm, T8 at 470mm, and Portland and T2 at 508mm). Independent feeders have stroke distances of 500-600mm (Table l-II) so are farther from the crust and hence plunger temperatures are lower. This will assist in keeping seals cool. As discovered in the cylinder trial, eccentricity is a major cause of rod seal failure (Section 7.3.3:5). Failed seals often look worn on one side in many of the plants the author visited. This is probably from insufficient side load support of the feeder and/or cylinder rod bush. The greater the angle of the rod the more the rod tears at the seals. Independent feeders can achieve more support and have less misalignment of the crustbreaker as the spool in not affected. This is a particular fault with the AEDD design. Portland development has been unsuccessful in providing a reliable bush low down in the dosing cup area as the bush designs tended to slow the spool and create jamming. Section 8.4.4 discusses eccentricity in more detail. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 295 Another factor on the temperature exposure of seals is the background temperature of the feeder in the superstructure. All plants surveyed (except Portland) have the ducting (that removes gases from the pots) at the base of the superstructure. Portland has the ducting at the top. This causes the heat to be drawn up around the cylinder and contributes to the seal cooking. Portland temperatures at the rod seals are about 155°C, with peaks of 330 C being attained for hot pots (Table 5-VII). Measurements at T6, T8 and T16 plants showed rod seal temperatures of 60-80°C; Atlas Copco, Parker, and Terry believe this is typical. Thus, Portland conditions are probably worse than most plants for temperature exposure on seals. 8.4.2:2 Normal Operation Because the plunger dips into the 950 C bath every 3 minutes or so, it was considered that reducing dwell time was an effective method of reducing temperature exposure at the seals. To find out how quickly the heat conducts up the plunger shaft, theoretical heat transfer calculations were made on infinite slabs, semi-infinite slabs, and irregular solids using formulae and tables from Holman (78). These showed that the temperature rise up the shaft was slow; minutes to hours were required before there was significant effect, rather than seconds as commonly thought. As this was hard to accept, a device was manufactured to test the temperature rise up a shaft. This was basically a Portland feeder minus the assembly. Thermocouples were placed at the top and halfway up the plunger shaft. It was the "worst case" design as there was no heat protection from alumina between the cylinder and the pot cavity. Some of the results have been noted in Sections 6.3 and 6.6 and shown in Figure 6-2. The figure shows that it takes over 50 minutes to raise the temperature at the seal area by only 50°C. This verifies the theoretical calculations. Clearly, plunger wet time (which is no more than 5s) has no significant effect on seal temperature. However, the seals reached the same temperature as the plunger in less than 16 hours. Therefore, plunger temperature is the most important factor in seal temperature exposure and, hence, seal life. It is interesting that almost identical results were achieved in a similar trial in 1993 by Patrick at Boyne Smelters Limited (83). CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 296 This conduction up the plunger shaft will cause both rod and piston seals to reach temperatures up to 300 C (Table 6-DI) which exceeds the maximum temperature rating of any of the normal materials used for seals (Table 8-II). This means that dagging is a probable sign that seal temperatures are excessive and seal life is dropping every second the dag is there. The dag is an indication of seals cooking. Control actions are the same as those discussed in Section 6.4. TABLE 8 -H MAXIMUM TEMPERATURE LIMITS FOR SEAL MATERIALS Material Operating Temperature Peak Temperature Viton 210°C 240°C Silicone 240°C 260°C Teflon 260°C 300°C (References 87, 88, 89.) 8.4.2:3 Hot Pots Plunger temperature surveys at Portland show a direct link between pot temperature and plunger temperature (Section 6.3.6:7). There is also a direct link between plunger temperature and seal temperature (Section 7.6.2:2). Hence, hot pot temperatures will affect seal life. (A "hot pot" is one with bath temperatures over about 970 C). Once a feeder is changed out in a hot pot, it is likely that the new one will also cook, leading to more changeouts. It is a good idea to avoid replacement until the pot is back to normal operating temperature. This procedure has been commonly used at Portland since mid 1993 with improved overall feeder life and less repeating feeder changeouts. Only when several feeders are leaking in the same pot are they changed. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 297 8.4.2:4 New Pots Protection of feeders from the extreme conditions encountered on start-up will benefit feeder life. New pot conditions may include: (a) pot temperatures over 970°C for extended periods, (b) high radiant heat from holes, and (c) high bath levels. Some plants put in old feeders with smaller plungers from off-line pots to assist the crust breaking and sticking problems associated with the larger anodes eg. Til and T17. This in effect protects the next overhauled feeder from heat exposure, and hence extends its life. If they worked prior to the scheduled pot outage, why overhaul them? Reusing feeders from old pots was trialed at Portland in April 1993, then reinstituted in February 1994, (after installation of spool inserts on all feeders) with savings of $12,000/month and no increase in failure rate. 8.4.2:5 Rod Seal/Bush Orientation To minimise the effect of high temperature on seals, it is best to have the seals mounted as close as possible to the cooler air of the internals of the cylinder (Figure 8-1). This idea also has the advantage that it gives greater protection of the seals from foreign objects/dust, as the bush is now between the wiper and the seals. Also the bush is closer to the driven mass, so there is better side support for uneven loading of the shaft when it is fully extended. This design cycles air about every 3 minutes, so the seal is continually being cooled. This is a very good design to help keep the seal cool. This design feature is used in Atlas Copco cylinders (Figure 1-20) and recent Terry designs for Portland, T6, T8 and Dl 1 at the suggestion of the author. Independent feeders are not affected by alumina ingress into cylinders because the alumina is quite some distance from the rod seal area (Figure 1-12). CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 298 8.4.3 Speed Based on best engineering judgement, slow speeds would seem to be better for seal life due to less frictional wear. Wear generally follows a cube relationship with velocity; for example, half the velocity relates to 8 times the life. The cylinder trials suggested a cubic relationship between speed and seal life (Section 7.3.3:4). Plant experience at Portland did not show this magnitude of improvement when the speed was changed in November 1991. This was due to a simultaneous drop in the amount of lubrication on seals (Section 7.5.4:2). This illustrates the importance of how one does an overhaul. Independent feeder plants and those with long stroke times have little problem with rod seal life; even for Viton seals which are of 50°C lower temperature rating to Teflon (87,88,89). This is partly due to lower frictional wear from the slower speeds (compared to integrated feeders). 8.4.4 Eccentricity If the piston rod does not move along its centre line, the eccentricity can cause more wear of the seal and leaks will develop. Useful actions to minimise the effect of eccentricity include (i) improving alignment by increasing bush length at the rod seal, (ii) maximising bush length of the shaft closer to the dosing unit and plunger, and (iii) using spigotted fittings at joints between assembly components. The integrated plant T2 had only a 2 year seal life versus 5-6 years for independent feeders with almost identical Atlas Copco cylinders (Table 1-H). If the life of cylinders are so different despite identical cylinder design, it is suggested that the different seal life is due to the feeder design and not to the cylinder. Better support for side loads would improve seal life. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 299 Support for side load can be improved at the cylinder itself by design of the bush at the rod seal (Figure 8-1). The Atlas Copco seal/bush length is about 50mm. Of the two types of cylinder used at Portland, there is little uneven wear and better life on Terry cylinders that have an 80mm seal/bush length compared to Parker which has a 50mm seal/bush length. Lindberg seal/bush length is about 50mm. A long seal/bush length helps alignment and hence seal life; the longer the bush the better the seal life. It may appear that there is little flexibility for plants that are restricted by cylinder length and are unwilling to spend capital on replacing all existing cylinders. One option is to split the bush in half and place the halves at each end of the front end cartridge. This effectively increases the bush lengthto th e full dimensional span of the front end cartridge. Thus the bush can be "lengthened" without changing the cylinder. Especially in a large cylinder with a long stroke, misalignment and side load can cause a very large force on the seals and rod bush, resulting in seal wear. Table 8-IH shows deflection for several plants. Note that independent crustbreaking units (D3, Dll) have only 5mm side deflection versus 20-30 for integrated feeders. Pechiney AP18 pots have a 500mm long cast iron bush housing (Figure 1-12) with (in the latest potlines) a 100mm long bronze bush at the lowest point. There is only a 1mm difference between shaft O.D. and bush ID. Alcoa feeders have no bush at the dosing unit except for the 6mm hardened insert invented for the Portland feeder (Figure 1-11). In addition, in the most recent Pechiney designs (eg. D10 Line 3), universal joints have also been installed to prevent misalignment. This is expected to result in longer feeder life by reduced eccentricity at the rod seal but at a very high cost. It is debatable if the cost (well over $100) is cost effective. It is desirable to have a bush as close as possible to the crust. However, for integrated feeders, the equipment at the lowest location is the dosing unit. The tighter the clearances between bush and shaft, the more the bush affects the dosing unit (spool) free movement which in turn affects shot size (refer Section 2.5.3). Thus, better cylinder life is constrained by the need to meet a primary duty of the feeder. Once again, the independent feeder is superior. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 300 TABLE 8-m PLUNGER SIDE DEFLECTION COLLAR TO % ROD SEAL AVERAGE EXTENDED CYLINDER TOTAL BOTTOM PLANT LEAKS AT FEEDER PLUNGER TO O.D. DEFLECTION CYLINDER WORKSHOP LIFE CYLINDER FLANGE BOTTOM FLANGE (mm) (mm) (months) (mm) (mm) T11 100 18 90 12 1041 1397 T9 100 19 70 24 - - T7 100 20 60 6 1043 1399 T17 100 20 75 36 1043 1399 T15 100 22 70 24 1092 1448 T16 100 22 10 24 1020 1376 T6 100 30 80 12 1022 1378 T13 100 33 90 9 1025 1381 T3 125 15 N/A 36 1260 1768 T10 125 15 80 N/A 1740 2096 T2 125 25 95 24 2555 3063 Portland 125 30 36 10 958 1466 D10 200 5 N/A 48 1000 1508 D3 150 12 N/A 48 1010 1518 Notes: (i) D10 have 70mm shaft in a 70.75mm ID cast iron bearing of 578mm long. This wears at the end, so another 120mm bronze bearing of similar ID has been installed. Deflection measured on 23.6.92. (ii) D10 line 3 incorporates a universal joint with its Atlas Copco 200mm cylinder. (iii) D3 have two 30mm x 60.5mm ID steel bearings to support the 60mm OD of four years such that the gap is shaft. Due to wear over a period now up to 4mm, they will be lengthened to 60mm and probably will be changed to bronze (based on the author's recommendation). (iv) T10 have same shape feeder as Portland but have a longer and smaller ID spool to support the shaft viz 1457mm not 832mm long and 52.5mm not 53.1mm respectively. The shaft is 50mm diameter. They were changing all feeders every 2 years whether they have failed or not until 1992 when they changed to failure based changeout after a visit by the author. (v) Atlas Copco recommend larger piston rod OD for integrated feeders to take side load. (vi) T13 rod seals mostly fail dueto electrical arcing, whereas other plants suffer from rod seal wear. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 301 At Portland, spigotted flanges were installed between the cylinder and the feeder assembly on 800 cylinders to improve alignment of the cylinder to the feeding assembly in 1992. No significant improvement resulted due to the fact that this was not the main cause of misalignment. The more important place to address is the bush at the feeder dosing cup which is closerto the crust to prevent twisting of the whole plunger shaft. A design was developed at Portland that effectively achieved maximum deflections of 6mm by use of bushes on or against the spool (76), but plant trials proved these seized in service due to clearances tightening under heat and alumina fouling the spool. It was concluded, therefore, that it is not practicalto install a bush against the spool on AEDD feeders without seriously affecting shot size. Another factor that can affect seal side load and piston rod breakages in AEDD feeders is the spool eccentricity and face. If the spool outer pipe is not exactly square, the spool will pull the shaft to one side and cause side load, stripping of the chrome coating from the piston rod, cushion bush failure and uneven wear of rod seal. Checking squareness in a lathe before installation is good practice. This is discussed in more detail in Section 2.5.8. Figure 2-17 shows the uneven wear of the spool prior to commencement of this procedure. 8.4.5 Rod Seal Material The recommended maximum operating temperatures of the three most common seal materials are shown in Table 8-II (87,88,89). Clearly, Teflon has a better temperature rating than Viton and has the advantage of having reduced friction. Comparison of the friction coefficient of different types of Teflon using the SRV from the lubrication trial is illustrated in Figure 8-2. It shows that carbon filled PTFE has half the friction coefficient of glass filled Teflon (used by the Alcoa USA smelters). CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 302 FIGURE 8 - 2 SEAL MATERIAL FRICTION COEFFICIENT 250 250 200 200 o o oz 150 150 O DC LL % 100 100 LL LL 8LU 50 50 CARBON BRONZE CERAMIC GLASS COMPONENT IMPREGNATED IN PTFE Note: SRV Analysis @50°C, 1hour, Steel roller - PTFE plate Anderson (Atlas Copco) states that Viton seals have a memory (90). They "remember" the hours of exposure at elevated temperature and when the hours exceed a certain critical signature time, the seals fail. Thus, a seal may be heated excessively over a period of time and not fail, yet for what appears no real reason, the seal fails after only a short period of high temperature. Anderson maintains that this explains why some seals survive days at high temperature and others fail in hours. What the observer is not aware of, is that the seals have gone through a period of high temperature on a previous period and are just about to fail when the short period of high temperature occurs. If this theory is true, then it is not just the temperature that is important, but the period too. Failure may occur in days at 150°C, hours at 200°C and seconds at 250 C. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 303 Portland, T2, T6, T8, TIO, D3 and smelters using Atlas Copco and CPOAC cylinders (Table l-II) used (or still use) Viton seals. If Viton seals are soft they can take more heat before they lose their flexibility. Parker seals possibly fail quicker than Atlas Copco due to Parker using harder seals. Thus, the method of manufacture, even for the same material, can have a significant effect on seal life. Most integrated feeders use Teflon. Based on the research of the author, since 1991, T6, T9, TIO and Portland have changed to Teflon, with T8 currently in the process of changing. Improved seal life has resulted in all plants. 8.4.6 Rod Seal Design The seal material is not the only requirement of a good seal. The shape of the seal, lubrication type (and quantity), preload pressure at the rubbing face at room temperature and the ability of the seals to cope with elevated temperatures are all important factors. A seal that does not leak at room temperature may be too tight at elevated temperatures and may erode at rapid rate due to the higher friction effect. This was found in development of the A3 dosing cylinder when seals that did not leak in the workshop seized the dosing unit at operating temperatures. Once the clearances were increased between the seals and the barrel/rod, the seizing disappeared. Seal design is a very complex skill. Trials have been conducted at Portland on many seal designs. Parker has trio led 2 types, Terry 3 types, Norton 1 type and Dover 1 type. In addition, all Parker cylinders have been converted to Terry front end cartridge units due to better relative life from Terry seal units. Figure 8-3 shows Weibull analyses for the three different types of seals used in Parker cylinders at Portland. One can see a 10 fold increase in life. The reversed seals (c) are benefited by better grease and less plunger penetration, so the effect is not all related to the seal design. However, the original Viton polypack and Terry front end cartridge are directly comparable as the same amount of grease was used. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 304 FIGURE 8-3 WEIBULL ANALYSES FOR PARKER ROD SEALS AVERAGE LIFE: (a) ORIGINAL PARKER POLYPACK (PRIOR TO NOV.91) 360 days 100 ^ 80 7 ,. 50 20 ^ '.'ir B = 1 .. (•^ •n =» ypjr Folll ire a 5 2 1 35 i lor 3d 1 1 .U 3 IDD. »&. GM, 1MQ. 9D0B. HMD. IO Total Opflrottng Tim* AVERAGE LIFE: (b) ADAPTED TERRY FEC (NOV,91-OCT,93) 1,800 days 100 ^L ^ 50 ::_^ _^r: I 20 5 B — 1 .' £ 5^ 2S P irj ire S *7 "C«r»i tor 3d 1 e :;s HI. 10Q. 20B- I MM. 2aai>- o ,7 Tatol Operating (Oays) AVERAGE LIFE: (c) TERRY REVERSED FEC (NOV,93-SEP,95) 3,600 days 100 50 20 soa. IODO- Notes: (i) Vertical axis is "cumulative percent occurred". (ii) "Censored" refersto replacement s that were not due to failures e.g. preventative maintenance. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 305 One observation in this research project and in the cylinder trials (Section 7.3.2) is the apparent fickle nature of seals. For example, Terry FECs (front end cartridge) work better in Parker than Terry cylinders despite beingAalmost identical rod seal cartridge. In fact, the Terry FEC in Parker has a shorter bush length which would result in less support on side load support (Section 8.4.4). Figure 8-3 can be compared to Figure 7- 5(a and b) to see the effect of the same seals in two different cylinders operating at the same speed. This shows that Terry seals had only 450-650 days life in Terry cylinders versus 1,800 days when the Terry seals were placed in Parker cylinders. The better results may be related to use of Teflon (rather than Viton) and better grease. The cylinder trials established that Teflon was superior (Section 7.3.3:2). Tranz 414 was used for Parker cylinders instead of Magnalube. The lubrication investigation established that Tranz 414 is better than Magnalube (Section 7.3). Norton and Dover have achieved 4-5 years life in plants that use Alcoa designed integrated feeders, yet when placed in Portland feeders, they were worse than the existing Terry design which achieves a one year life. This suggests that Portland conditions are worse than other Alcoa locations. It is important that in any trial of alternative seals in an air cylinder, all practical conditions are identical before comparing units. Til, T13, T16 and D12 have double rod seals in their cylinders. These are made by Norton or Dover. At Portland, Terry have double seals and Parker have single seals. Atlas Copco and CPOAC have one seal. T17 use a single Dover seal with a four year average life. If the cause of failure is exposure to high temperatures, it is most likely that all seals will fail with a similar life as they are all in contact with the same hot surfaces. In the case of the plants listed above (viz. T6, T9, T10, Portland, Til, T13, T16, D12 and T17), the best seal life is at T17 which has single seals. All the plants have identical bore and stroke cylinders using identical assemblies. Hence, it is concluded that multiple seals are probably not worth the extra expense. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 306 There are two common ways to keep a pressure behind the seals to ensure they press against the piston rod. Most commonly an 'O' ring of silicone or rubber is used, but some manufacturers use stainless steel spring type energisers. As noted in Table 8-II, silicone has a higher maximum operating temperature between Viton and Teflon. If temperatures are high enough to affect Teflon seals, it is likely that silicone energisers will fail before the Teflon does. "O" rings may flatten due to thermoset, thus reducing seal pressure or allowing air leakage between the seals and the cylinder casing. This occurred for several runs in the cylinder trials (Table 7-1). If 'O' rings have to be used, rectangular energisers appearto b e better than circular ones as there is more even wear at the seal surface due to the pressure being applied over a greater area. Use of stainless steel is becoming more common for Teflon seals, as the weakest point in the design of most seals has been the 'O' ring type of energiser. These are used by Norton at T17, and recent Parker designs at T8 and trial units at Portland. This design is recommended as it totally prevents energiser failure and the pressure of the seal against the rod is not affected by thermal expansion which is a problem with Teflon seals. 8.4.7 Rod Wiper Design The wiper in the front end cartridge is designed to keep alumina out of the cylinder in the seal area (Figure 1-20). Portland and T8 use bronze wipers on the rod seal cartridge and both suffer from grease breakdown. Shell advised in the lubricant investigation that copper and iron are catalysts to the breakdown of grease (Section 7.4.4:3). In October 1993, Portland wipers were changed to Teflon to avoid both the pitting and grease breakdown. Since this time, seal life has deteriorated, although there may be insufficient data to date to confirm a distinct relationship (Figure 8-2). CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 307 8.4.8 Venting Rod Seal Air Leaks As noted in Section 8.3.1, rod seal leaks can aerate the alumina in the dosing unit or overpressure the cavity where the feeder is mounted due to inadequate venting of the leaking air. It is preferable to design a large vent pipe (that is angled downwards) to prevent alumina from collecting and blocking the vent. This pipe is best vented into the pot rather than out of the superstructure to avoid alumina spillage on a bad air leak. Routine air leak checks on each pot will ensure that bad air leaks are not overlooked. Section 9.4.2 explains how this is done. One plant (T9) which has little problem with ore leaks has a snorkel type vent pipe so that the air outlet is above the alumina level in the superstructure (76). This is a very good idea and simple and cheap to fit on pots that use overhead cranes to fill superstructures. One plant with good ore leak history (T13) suffers from loose mounting bolts. Although T15 shows a high rate of failure for ore leaks (80%) this is mainly from no lubrication on Viton seals, but they also have loose flanges that not only cause ore leaks but also arcing as the feeder moves sideways and touches the oppositely charged superstructure. Most other integrated feeder plants that have better success with ore leaks have an adequate vent pipe and/or better seal quality. The worst plant seen was at T9 where Hanna rod seal leaks were so bad that they could not even hold the plunger retracted. These have now been changed to Lindberg. A design developed as part of this research is the fixed installation of insulation washers on the top flange with a gap for air to vent from the feeder cavity (Figure 8-4 and 8-5). The boomerang bracket design not only solves the air vent problem but also has the advantages of less parts (6 parts instead of 25 on the old design), is quicker to install (less time people are exposed to heat, dust, and fumes) and prevents arcing (refer Section 8.7.1). The air vent hole is equivalentto a 75mm O.D. vent, so is far superior to any of the vents commonly installed in pot superstructures. The design costs only about $40to install and is easy to retrofit if plants currently use insulated flanges. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 308 FIGURE 8 - 4 BOOMERANG BRACKETS AND FLANGE INSULATION INSULATION AT BOOMERANG BOLT HOLES BRACKETS CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 309 FIGURE 8 - 5 MOUNTING BOLT AND BRACKET AIR GAP SIDE VIEW DETAIL X MOUNTING BOLT BOOMERANG BRACKET BOOMERANG BRACKET VIEW A STEPPED WASHER MOUNTING (ATTACHED TO BOLT) BOLT FEEDER NSULATION MOUNTING NEMA GRADE G1 1 PLATE SUPERSTRUCTURE FLANGE BOOMERANG BRACKET DETAIL X CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 310 8.5 PISTON SEALS Until recent years, most integrated feeder plants did not realise piston seals were a problem area. Following tabling of the author's "Feeder Recommendations" in 1992, all Alcoa plants investigated piston seal leaks (Table l-VQ). In almost all plants, piston seal leaks proved to be a major cause of air consumption and, in many cases, it also severely affected feeder operation. Dll reduced air usage by one compressor by addressing air leaks. T6 found air leaks were a major contributor to anode effects. The comments in Section 8.4.5 on rod seal material and temperature exposure also apply to piston seals as the conditions are similar. Pistons usually have one or two seals and one or two bush rings to keep the piston running parallel to the centre line of the cylinder. It is possible to have a seal leak in one direction only, so it is important that two air leak tests are conducted with the cylinders energised in both directions, otherwise leaks may not be detected. In the A3 trials, seizing of the piston seals occurred due to clearances being too tight (Section 10.6.3). A small air leak at room temperature is probably satisfactory for piston seals to ensure they are not too tight at elevated temperatures (Section 8.4). No significant trials on alternative piston seals were conducted as part of this research due to its lower priority in integrated feeder failures. It is expected that further research will be needed in the future as the life of other components increase. 8.6 BROKEN PISTON RODS 8.6.1 Introduction Most plants surveyed indicated fatigue failures such as broken piston rods and broken spools in (Table 1-V). Usually less than 5% of overhauls were done for these reasons. T17 had 25% piston rod failures with Portland and T6 both having 10% piston rod failures. Portland also had 30% of spools fatiguing in 1990 probably due to excessive speeds of the plunger. No sign of piston rod failures has been reported for independent feeders presumably due to the lower piston speeds. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 311 The problem of fatigue was tackled by (i) finding the stroke time to break a hole and (ii) finding the minimum stroke time to avoid mechanical failure. Once these were established, either the stroke time had to be reduced to minimise fatigue, or continue with existing failure rates. Ultimately the hole has to be broken and, if this resulted in failure of the rod, the costs would have to be accepted. Trials of timeto break a hole at Portland showed that one could do so at zero velocity on a continual basis (Table 3-HI). Thus, there was considerable scope to reduce the speed of the plunger. Hence, it was decided to investigate slowing the piston speed to achieve acceptable fatigue levels. 8.6.2 Fatigue Analysis Dr Hugh Stark (Unisearch, University of New South Wales), an expert in fatigue, was engaged to establish what speeds resulted in fatigue (86,91,92). He conducted a series of tests of 14 piston rods from the two Portland cylinder types using an Instrom 8504 load tester at various stress loads from 13-60 kN at 10-20 Hz (86). Once the limits of the piston rod were established, one needed to find out what stroke speed was necessary to fall below the critical stress level. This was then cross checked against the stroke speed necessary to prevent fatigue of the spools. Similar tests also found that the stroke time to prevent fatigue of spools was similar to that for piston rods (Section 8.8.1:1). Thus a final setting was established to minimise all fatigue. The fatigue analysis showed that Parker 10L45 piston rods were not as good as the Terry piston rods made by Kempe Engineering from 4140 induction hardened steel. Failure was partly related to the design of the run-out of the piston rod thread as well as the type of rod material. There were basically three different types of failure as illustrated in Figure 8-6. The Terry breakage occurred at higher stress levels and with a less crescent shaped profile on the break surface. The latter feature suggested a less brittle rod material than the Parker rods. These types of failure are similar to that found at T6 and T17. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 312 FIGURE 8 - 6 SKETCH OF BROKEN PISTON RODS BREAK TYPE1 TERRY POINT ACROSS FIRST THREAD TYPE 2 TERRY BREAK POINT 1-3 THREADS INTO MASS 1£*A3! BREAK TYPE 3 PARKER POINT ACROSS FIRST THREAD CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 313 A random sample of 5 Parker and 5 Terry cylinders were tested at the Portland workshop at various stroke times to find their shaft deceleration rates under varying speeds. Using the test data from Unisearch, it would be expected that about 40% of Terry and 60% of Parker rods would fail. This closely followed plant experience with Terry piston rods. The surprise from the tests was that Parker rods had worse results despite a lower failure rate in the plant at that time. Weibull analyses also showed similar results to Unisearch. About 6 months later there was a sudden increase in failures of Parker rods. The reason why the failures were not high at the time of the Unisearch tests was that the average life of rods was low and they had not reached their critical failure life. One could say it was the "honeymoon period". This experience substantiates predictive methods as a worthwhile tool. There was little change in average deceleration rates once a stroke time of 1.5s was achieved as at this point the cushioning on the cylinder became the dominant factor in speed reduction. The drop in stress was by about a factor of 3 for a change from 0.8s (which was the existing stroke time) to 1.5s. These data suggested that a stroke time of about 1.5s would prevent fatigue. Note also that testing at Portland showed a higher deceleration rate going up. than down viz. 800ms'2 versus 500ms"2. This is partly due to the collar hitting the cup before the spear on the top of the piston engaging the top cushion viz. 50mm versus 25mm from the top respectively (Figure 1-11 and 1-20). The peak upward velocity from plant data was l.Om/s (Table 3-1). At this speed the spool will close a 2mm gap between the stationary assembly housing and the moving spool cup in 0.002s. The ricochet causes the high deceleration rate on the shaft. In a number of plants visited (T9, Til, T13) there was evidence that there was in fact no gap at all between the cup and the stationary feeder assembly. This metal-to-metal contact can lead to fatigue of the feeder components and explains their history of broken rods, spool failures and failure of piston/piston rod connections. This can be a problem in independent feeder dosing unit where the spools almost (or definitely do) touch the housing. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 314 8.6.3 Results and Conclusions of Fatigue Tests Based on the above discussion, the stroke time of Portland feeders was increased in November 1991 to 1.6s (downstroke) and 1.9s (upstroke) from original stroke times of 0.9 and 0.8s respectively to reduce the speed and hence kinetic energy of the moving components of the feeder. The target of 1.5s downstroke was not quite achieved and a smaller upstroke was chosen as this was not a critical factor for shot size control. The results of the change in piston speed is seen in Figure 4-5 (for plant changeout rates) and Figure 8-7 using Weibull predictions over time. A reduction of about 70% in piston rod failures occurred in the plant and the Weibull analysis showed an increase in average life from 1,500 days to 2,600 days. In addition, the thread run out was minimised and all new piston rods were changedto stee l grade 4140 for better fracture toughness. The latter change had little effect. FIGURE 8 - 7 WEIBULL ANALYSIS BY QUARTER OF PISTON RODS WEIBULL BY QTR REPORT from 1 January 19H9 tD 3D September 1 395 Repair Part 421 (PISTON &0D). 521 (PISTON ROC) 1 2 5O0O 100 N go u L 1 4O00 \/.V \ so? f 3000 Ill 1 so V n 2000 40 T 0 30 I a u 1OO0- 10 r IO « 0{T- i r p*"' I ' ' I ' M'' T ' T"l " I ' ' 1'» ' I ' 'I'M"' • i \ | i i i1 i 'I |~ JAJ0JAJOJAJ0JAJ0JAJ0JAJOJAJ APUCAFllCAPUCAPUCAPUCAPUCAPU MRLTMRLTMRLTMRLTNRLTMR u T tj 5 k B8a89999999&999999999S999ft 99390000 1 1 1 122 2 23333444-4-555 Month Re paired ,m.= 20K **6=50K **# = 80« - = Number Failures Months wlffi leas rhan 5 failures Ignored Date Created: 2 October 1995 Values aver 500D days truncated 1 SLOW SPEEDS 2 DUAL DWELL TIME CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 315 Following advice from Portland on piston rod design, T17 changed the piston design and have reduced failures by 75%. In addition, trials of smaller DCVs indicated A that slower rates have not affected hole breakage and longer piston rod life is expected. The trial of smaller valves was expanded in 1994 at T17 to more pots. Hence, plant results verified the investigation on fatigue of piston rods. The results of this study also questioned accepted theories, some 30 years old, that kinetic energy is requiredto break feeder holes (Chapter 3). 8.7 ARCING j\rcing occurs if there is an electrical path between the crustbreaker and superstructure when a pot is on anode effect. Insulation is generally achieved by installation of high impedance washers and rings between the top feeder flange and the superstructure and similar material at the bottom of the feeder assembly. Some plants also specify their crustbreaker cylinders to be insulated. For the 11 integrated feeder plants surveyed, arcing occurred in half of them. (Figure 1- V). There have been no reports of arcing occurring on independent feeders due to better insulation and less time in the bath, although the AP30 has better insulation than the API8 to prevent electrolysis. 8.7.1 Insulation of the Assembly Plants with a history of arcing usually have problems with loose mounting bolts which allows the feeder to move sideways and reduces the clearance between feeder and superstructure. This study identified the cause of loose mounting bolts was mechanical failure of the insulation at the feeder-to-superstructure mounting flange. Use of NEMA grade Gl 1 (at least) insulation is recommended. In addition, it is best to use large steel washers to avoid excessive point compression that will break down the insulation. Failure of feeders due to loose mounting reduced from 30/month to under 7/month after changing to Gl 1 insulation (from G9) and using bigger washers (Figure 8-8). VO Hi ra z •s z in w Z •4 p ^£ ro a 2 £ <"S ra x S (a tCHL a"- w m LU Z .-1 H z |Utl»3 ai w a w w o XS w oao CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 317 This further dropped to under 2/month when the mounting flange was changed as discussed in Section 8.4.8 and shown in Figures 8-4 and 8-5. This design could be at many smelters. It is cheap, simple to install and provides adequate surface to protect insulation from mechanical damage as well as other advantages to feeder failures and exposure of people to heat, dust and fumes. At Portland, the bottom feeder assembly insulation cannot withstand the temperature - peaks of over 400°C are evidenced. The $10 silicone rubber insulation bands fail within days of installation, but, as long as the feeders are not moved, the bands continue to provide adequate protection even when mechanically broken. Many alternative materials were trialed but these were either too expensive (some over $100 each), could not be retrofitted easily, or were mechanically weak. Those trialed included Teflons, ceramics and ceramic sprays. It was finally concluded that -fc*s was the cheapest solution. The insulation used by Pechiney, T2 and D3 cost several hundred dollars but are almost never changed. The Portland design is a most cost effective solution, especially if the feeder has a lower average life. 8.7.2 Cylinder Insulation Alcoa, Alcan and several other companies specify that the piston rod and piston should be electrically insulated from the stationary barrel and end blocks of the crustbreak cylinders. Insulation costs about 30% of the cost of a new cylinder. Most cylinders from Parker, Atlas Copco, and CPOAC are not insulated. Some plants that claim to have insulated cylinders (and specify this requirement in the purchase of new cylinders) generally have stainless or bronze cartridges (T17, T8), graphite impregnated seals (T8), or have an electrical path between block, guide, spool and shaft (all AEDD feeders). At most plants using AEDD feeders (T6, T9, Tl 1, T13, T15, T16, T17), the feeder is virtually touching the superstructure at the bottom of the feeder. If the feeder is not placed exactly central on the mounting frame of the superstructure, there is an electrical path. In more recent AEDD designs (Portland, TIO, T7, T4), there is an insulating ring at the rim area of the feeder (Figure 1-11). CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 318 In the plants that have the worst arcing problems (T13, T9, T15), the problem is aggravated by poor mounting bolt securing. Loosening of the feeder bolts allows the feeder to move sideways so the feeder touches the assembly at the feeder base. Mounting of the feeder and the design of the feeder itself are the main causes of arcing, not the cylinder insulation. So insulating the cylinder is pointless; a waste of money. Tests were conducted at Portland on 34 cylinders when they arrived in the workshop for overhaul. Nearly every one monitored showed no insulation capacity. After routine overhaul with new seals, bushes and wipers, some 75% had more than 10 mega Ohm resistance at 1,000V DC. One can assume that this sample is typical of the population, and that insulation breaks down in service. This was verified in the Portland/Parker cylinder trials where all cylinders tested had lost insulation after rod seal failures despite using insulating grease (Section 7.3.3:12). In most cases, plants that already have insulated cylinders would find it expensive to take the insulation out. Mostly the insulation does not fail mechanically so there is no maintenance cost in continuing to use it. However, it is recommended that, if there is a need to purchase new cylinders, it is unnecessary to specify that the cylinder must be insulated; thus saving about 30% on the cost of the new cylinders. Note too that not specifying insulation in cylinders opens up opportunities to use seals that may be based on graphite or other conducting materials. The friction coefficient of graphitefilled Teflo n is lower than other Teflon based material so may provide longer life (Figure 8-2). 8.7.3 Other Methods to Control Arcing Feeding on anode effect increases the time plungers are in the liquid which in turn increases the magnitude and frequency of high voltages between feeder and superstructure. Whatever the quality of the insulation, the least the plunger enters the liquid and the lower the voltage of the pot when this occurs, the lower the probability of arcing. As mentioned in Section 5.4.5, there are advantages to plunger life if the plunger does not cycle into the bath when the pot is on anode effect. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 319 There are distinct advantagesto reduction of the chance of arcing it the plunger does not touch the bath at all. Clearly, if no circuit is made there can be no arcing. Choosing low plunger penetration may not always mean that the plunger does not touch the bath, but it at least reduces the frequency. As discussed in Section 5.4.2, plunger life is far superior if it does not touch the bath or at least has minimal penetration as corrosion is one of the dominant causes of plunger wear. One other option to prevent arcing is to insulate the plunger from the plunger shaft (as done at T13). This breaks the circuit. This is a costly solution, restricts the use of standard plunger materials and cannot be used with electrical end stroke sensing. For older plants with very tight clearances between feeder assembly and superstructure this option may be solution, but one should consider a more permanent solution such as better mounting flange insulation washers or reconsider the strategy of feeding on anode effects. 8.8 DOSING UNIT FAILURES 8.8.1 Broken and Worn Spools 8.8.1:1 Integrated Feeders Spools on Alcoa AEDD feeders fail by two mechanisms; wear and fatigue. Wear is considered in Section 8.4.4, where the effect of eccentricity on rod seal wear was discussed. This problem can be solved by use of a bush on the inside bottom of the spool and/or by checking clearances in the assembly when the spool is manufactured. Portland has had little to no wear since these checks were initiated in 1992. Spools of Alcoa feeders suffer wear from misalignment and inaccuracy of manufacture. Checks on 60 spools at Portland in 1991 showed that 90% of spools had wear (Figure 2-17 and Section 2.5.8). A similar trend was seen by the author when visiting T2, T6, TIO, Tl 1, T13 and T15 in 1992. It is good practice to turn the spool in a lathe prior to use to ensure all surfaces and edges are symmetric. It takes a few minutes but it may save the cost of another spool and/or bad shot size. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 320 Portland installed a hardened insert from 1990 to prevent spool wear (Figure 1-11). This was later changed to bronze in 1992 as the plunger shafts were wearing. There has been minimal wear on the spool since "hardened" inserts were installed. This design can be installed in almost all AEDD feeders. Fatigue of spools on integrated feeders generally occurs by a horizontal fracture at the top of the dosing unit at Portland. Consideration of the spool fatigue failures was carried out by use of computer stress analysis and laboratory tests by Dr. Hugh Stark (91,92). Two actions were taken to reduce fatigue; slower speeds and vertical webbed sections to reinforce the horizontal plate of the spool. Stark found that both these factors extended spool life. Testing of deceleration rates of the spool showed that there was a marked drop of deceleration rate for 1.5-2.0s stroke time (as was found for piston rods in Section 8.6). Webbing was introduced in July 1991 and feeder speeds were reduced on all pots by November 1991. Although the target was a stroke time of 2.0s, the downstroke was a little quicker at 1.5s. The result was a median life of about 10,000 days. This relates to virtual infinite life. Hence, the changes achieved the objectives. 8.8.1:2 Independent Feeders Because the independent feeder spool cylinder is so small there is less speed, however, the cushioning can be inferior. D3 has had problems in this area. Because the independent feeder spools generally have very small clearances between spool and fixed seat, any error in stroke length on assembly can result in metal-to-metal contact which can cause a considerable shock to the spool shaft fittings. It is very important that there is a clearance between these faces but not large enough to cause an alumina leak. One must also consider expansion differences between room temperature assembling and pot operation. A tight clearance at room temperature may be much tighter at elevated temperatures which could lead to fatigue CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 321 8.8.2 Spool Jamming One cause of failure that occurs on integrated feeders is "not feeding". This is often caused by complete jamming of the spool. Partial jamming of the spool can result in inconsistent dosing of alumina that can often go undetected. Ore leaks can occur if the spool does not seat fully after feeding the shot. Jamming can occur because of foreign objects fouling the spool or by incorrect spool manufacture. This problem was addressed at Portland by: (i) better installation techniques (explained further in Section 9.6.1:2), (ii) increasing insulation rating from NEMA Grade G9 to Gl 1 (Section 8.7.1), (iii) drilling holes in the assembly (Figure 1-11), and (iii) changing the mounting flange (Section 8.4.8). Prior to mid 1990, 25% of Portland feeders failed from jammed objects (which includes spool jamming). This has now droppedto almos t zero. Jammed objects at most plants are generally nuts, bolts and washers from the mounting flange falling into the hopper. The reasons the bolts come undone in service is related to insulation quality, steel washer size being too small (causing the insulation to squash), poor cushioning or too small an insulation washer. It is also important that no foreign objects are let fall into the alumina hopper. For most integrated feeders, the spool is the dosing mechanism in the integrated mechanism (Figure 1-8). For most independent feeders, the dosing unit is a smaller device which is driven by a small (40-50mm OD) pneumatic cylinder (Figure 1-2). Both can suffer from similar problems. Big improvements have been made at Portland on its AEDD feeder when holes were drilled in the superstructure into the spring chamber (Figure 1-11). The reason for drilling holes in the assembly was to prevent a partial vacuum occurring in the spring CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 322 chamber. This vacuum tended to drag alumina into the gap between the spool and the assembly and lead to jamming. Holes and cutting the spool guide solved this problem by letting the chamber breathe (Section 2.5.8). This problem with alumina being trapped between the spool and the assembly is probably present at all AEDD plants but is probably not detected as most plants do not do any shot size testing on operating pots and the fact that feeders are cleaned before they get to the repairers. There are few repair technicians who will not relate stories of feeders with no faults being returned to the workshop. These are probably due to alumina jamming the spool. This modification is recommended in all AEDD plants. Jamming can occur on the 150mm outer pipe of the AEDD spool (Figure 1-11) or the cone of the independent feeder because the spool: (i) is oversize - too tight a clearance to the assembly allowing alumina to pack in to the gap and seize. (ii) has a bottom face not horizontal - this causes the spool/cone to rotate vertically and foul the shaft and the outer assembly casing. (iii) is not central - if the outer pipe/cone is not exactly on the centre line it will foul one side. (iv) has an outer face not parallel to centre line - this causes the spool/cone to rotate vertically and foul the shaft and outer assembly. Specific data on dosing unit failure of independent feeder plants was not available in this study. However, it is likely that the above factors could be a problem with any type of dosing unit unless the quality control of manufacture is very good. As independent feeders have smaller dosing spools, the faults are expected to be less than for integrated feeders. Many of the problems of quality control of manufacture may be minimised in independent feeders due to the use of pneumatic cylinders rather than springs. As discussed in Section 8.8.3, there is a much higher pressure for cylinders, so many errors can be overcome by more force. Once again, this shows the robustness of the independent feeder designs. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 323 Note also that most plants do not check the accuracy of their feeders. Hence, they may well have a problem with their feeder unit accuracy and do not know it. 8.8.3 Spring Failure This type of failure applies to AEDD integrated feeder designs only, as independent feeders do not have springs. Note that the Alcoa specification for AEDD feeder springs specifies 302SS not 304SS which is more readily available. The former is a better spring material, but only marginally so. It is unclear why AEDD feeders have stainless springs as stainless is a very poor material for springs as creep is possible. It could have been due to prevent magnetic effects, but one usually finds that stainless steels gradually lose their non-magnetic properties after a period of time in the potrooms, so this is not relevant. Whatever the reasoning, it has been found that steel springs work fine in Portland and T16. Other plants are now changing over to steel springs from stainless springs. They are cheaper and work better. T8 (an integrated feeder but not an AEDD) uses the stroke of the plunger shaft to control the dose and does not have a spring to control the dosing spool. The shaft itself is the spool, thus reducing the feeding components and potentially increasing life. However, the seal between the shaft and the fixed assembly has a vertical rather than a horizontal sealing face when extended, so there is more opportunity for alumina bypass. Spring failures are reported at several plants using AEDD feeders. Springs fail by reduction in length generally with a colour change. The shorter the spring the darker the colour. This problem is due to creep of stainless steel at elevated temperatures; over 300 C. The higher the temperature, the darker the colour. The effects of spring length change include: (i) slower spool movement causing alumina bypass and affecting shot size variability (refer Section 2.5.3); (ii) less closing pressure on the spool causing leakage when in the down position and hence larger dose size (or leakage during the complete dwell time); CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 324 (iii) less cushioning of the spool when the shaft reaches its top position, contributing to fatigue (not normally a major problem in most plants). Hans Kempe (Kempe Engineering) has designed a high temperature die steel spring that provides a larger force which, in tension, is not affected up to 500°C. Figure 2-12 compares spring tension of 304SS and Kempe springs. This shows how the stainless spring used at nearly all AEDD plants (until this investigation) shrinks with temperature. Portland, T9, T15 and T16 are in the process of changing over to high temperature steel springs on a needs basis on overhaul. If one compares the force of springs used on Alcoa feeders with the cylinders used on independent feeders, the independent design is far superior (Table 8-IV). TABLE 8 - TV AEDD SPRING FORCES Stainless steel spring 230kN Kempe high temperature spring 420kN Air cylinder 1,350kN Basis: springs at 200°C fully extended (Figure 2-12) versus 50mm O.D. cylinder at 700kPa. Figure 8-9 shows the Weibull analysis of stainless (part number 312) and Kempe high temperature springs (part number 321). A Weibull analysis is a log-log Y axis of percentage failed, versus a log X axis of life. The analysis clearly shows that Kempe springs are superior to stainless steel springs with a 50% failure rate of 3,500 days versus 1,800 days. T16 started to install steel springs of another design in 1992 and similarly found that they performed superior to stainless steel. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 325 FIGURE 8 - 9 WEIBULL ANALYSIS OF SPRINGS AVERAGE LIFE: 1800 days Repair Part 512 (SPRING (STAINLESS)) 100 ~y. y' - ,s 80 y y -y S 50 y y *' 20 10 ft — i -i 22 1 J'f ^ •i 1 FOTIL s 1 3C 3 = s 97 :r tor 3d 10- 60. 10D. 200. fiOCL 1000L soon. 1000a. Total Op«rorTr»g Time (Cays) AVERAGE LIFE: 3500 days R^pflirP-art 521 (.HIGH TEMP SPRING J 100 80 ..£ _-""-- - 50 h- > . ^ ^y" 20 4* 3" 10 4 ,ylf? V ft = 1 .1> f.+ T} =S 52 -^ ,> •r*- * re'i • 1 54 tor ad •• 8£ • 10. 20. BO. 1DD. J»B. SML 1WWL aUOBL 6WB. I*HKM. Tatol Op«rorTr»g TIFT» (Dajre) Notes: (i) Vertical axis is "cumulative percent occurred". 00 "Censored" refers to replacements that were not due to failures e.g. preventative maintenance. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 326 8.9 MAJOR FINDINGS OF EXTENDING FEEDER UNIT LIFE (i) Attention to piston thread runout and piston speed reduces fatigue of piston rods significantly. (iii) Lengthening cylinder rod bush extends seal life by protection from side loads. (iii) Reusing feeders from off-line pots extends feeder life. (iv) Having the rod seal closest to the cylinder cavity reduces temperature exposure of the seals. (v) It is just as important to optimise energisers behind the seals as the seals themselves. Stainless steel is better than silicone or Viton. (vi) In theory, Teflon rod wipers are superior to bronze as bronze wipers catalyse grease and do not wipe the rod as well as Teflon. (vii) It is not necessary to insulate cylinders if the crustbreaker/feeder unit is insulated from the superstructure. (viii) Cylinder insulation breaks down with time. (ix) Spool fatigue on integrated feeders can be reduced by webbing and slower speeds. (x) Attention to spool manufacture quality control will reduce wear and poor shot size accuracy. (xi) Attention to mounting bolts, insulation and foreign objects reduces fouling of the dosing system of feeders. (xii) High temperature springs are superior to stainless steel springs, but not as good as air cylinders for dosing spool movement. 8.10 RECOMMENDATIONS OF EXTENDING FEEDER UNIT LIFE (i) Minimise crustbreaker speed to reduce fatigue. (ii) Reuse feeders from pots that are scheduled to be taken off-line for startup pots. (iii) Locate rod seals next to the cylinder air cavity. (iv) Do not insulate cylinders. CHAPTER 8 EXTENDING FEEDER UNIT LIFE PAGE 327 (v) Feeder spoolsto b e made to drawings. (vi) Ensure mounting bolts are secure and insulation is evenly loaded. (vii) For plants that have space, install boomerang brackets on the feeder mountin: flange. (viii) Use high temperature springs on AEDD integrated feeders. CHAPTER 9 COST REDUCTION PAGE 328 CHAPTER 9 COST REDUCTION 9.1 SUMMARY Statistical techniques have been used to analyze data on failure rates of parts and components on feeder/crustbreaker units. These showed opportunities for maintenance cost savings. Non-traditional methods of plunger design using bimetal, high nickel alloys and short length can reduce costs by up to 90% with corresponding increase in life by up to 15 times that of cast iron. Complete changeout of all parts every time a feeder gets to the workshop leads to excessive costs with no improvement in life. A repair strategy based on air leak tests can reduce costs by up to 30%. Tracking systems are not only useful for day to day control, but are critical in assessing the effective life of any design changes and improving design in a cost effective manner. CHAPTER 9 COST REDUCTION PAGE 329 9.2 INTRODUCTION Historically, on failure of crustbreaker or feeder units, smelters have routinely changed out the plunger, all cylinder seals and any other items that were detected as being faulty. Costs averaged about $500-700 (Table 1-IV). This Chapter questions this practice as current maintenance strategies for maximum reliability suggest a more statistical approach is more cost effective. Use of Weibull analyses, control charts and Pareto charts along with an effective tracking system can identify what items need to be changed based on their current life and the failure history of each part. It is also important if one wants to reduce failures, to link the reason for removal with what the tradesperson finds in the workshop. This has often not been researched, such that modifications and design activities have concentrated on areas that are not necessarily the highest risk or greatest failure items. A tracking system that links: (i) where the feeder was installed, (ii) reasons for removal ("operator observation") and (iii) what the repairer sees ("mechanical faults") is critical to improving feeder life. This Chapter draws on different attitudes to the traditional method of feeder maintenance. Main areas examined are. the use of exotic materials for plungers, use of plungers made from two different metals ("bimetal" plungers), acceptance of some cylinder air leaks at maintenance, acceptance of air leaks on the pot, not overhauling feeders from pots that are taken off-line, uses of an effective tracking system and quality control both into and out of the workshop. Irrespective of the repair philosophy used, the best method to reduce feeder costs is to reduce the number of feeders that reach the workshop; install parts that will last the longest time almost irrespective of cost. About 75% of repair costs are for parts, so one CHAPTER 9 COST REDUCTION PAGE 330 must maximize the performance of the critical parts, rather than concentrate too much on labour costs. 9.3 PLUNGER SAVINGS In Chapter 5, the causes of plunger wear were identified and it was explained why different materials were trialed and their relative success stories (Table 5-IX and Figures 5-12 and 5-13). The relative wear rates in comparison trials indicated the materials and designs that achieved the longest life. However, the longest life plunger may not necessarily be the most economic. Depending on the plant concerned, economics may outweigh the life prediction, especially if one tells accountants or managers that the cost of new plungers may change from $40 to $400. Financial considerations have historically restricted usage of exotic materials as most engineers had little to no evidence to justify the more costly materials. To consider economics of different plunger designs, one needs to consider the plunger length (or penetration depth) and whether the complete plunger is made of the same material. From discussion of the effect of penetration depth, Portland has tried two different plunger lengths (255mm and 305mm) which result in 50mm different plunger depths; 305mm being the deepest. To minimize the cost of plunger tips, a "bimetal" design was developed where the lower half (which gets wet) is made of expensive passive material, and the top half (which does not get wet) made of inexpensive steel (Figure 9-1). This design was developed with significant assistance from Hans Kempe who has considerable knowledge on metallurgy and thread designs for arduous environments. The bimetal design almost halves the cost of stainless plungers. For example, Nicrofer price drops from $480 to $300 for 305mm plungers (Table 5-TX). The bimetal design is less CHAPTER 9 COST REDUCTION PAGE 331 FIGURE 9 - 1 BIMETAL PLUNGER PLUNGER BODY MILD STEEL ANTI SEIZE THREAD PLUNGER TIP:HIGH TEMPERATURE CORROSION RESISTANT ALLOY CHAPTER 9 COST REDUCTION PAGE 332 economic for cast tips because of the expensive drilling costs. For example, Niresist complete plunger costs $165 against a bimetal cost of $145 (Table 5-TX) with a steel adapter of $90. Table 9-1 shows detailed calculations on the cost per tip/day for 14 plunger designs of the best 11 of the 19 plunger materials tested at Portland for different plunger length and complete/bimetal designs. Details are given using "within-pot" trials (where comparisons were conducted in the same pot for the same time at similar starting diameter) and plant data directly from workshop measurements at overhaul. "Within- pot" data is far more representative of relative wear rates than plant data as the variability is much lower due to no pot-to-pot differences. However, plant data is a useful cross-check on limited trials. In general, the trends are similar for "within-pot" and "workshop" data, but the plant data is more conservative with respect to life. Three bases have been considered for the calculations of cost/tip day. (i) Material cost only: The cost of manufacture of a plunger is divided by the life. This price is conservative as there are handling costs in overhauling a feeder, even if the only problem is the plunger. Usually there are many other problems that exist when a pointed feeder is removed, as usually the feeder has been in operation for a very long time. Most plants tend to do a complete overhaul for pointed feeders to avoid a quick return into the shop for some other reason. (ii) Labour cost only: This assumes that for every plunger failure, $75 is incurred due to transport, overhaul checks and replacement of some basic minor parts. For plants which use contract labour for feeder removal and replacement (e.g. T6) the cost will be much higher. This compares with an average overhaul cost of $635 for Portland and $718 for Alcoa in 1991 (Table 1- TV). At Portland, in 1994/95, about 30% of feeders pass through the workshop without major work at a handling cost of $75 each. CHAPTER 9 COST REDUCTION PAGE 333 TABLE 9-1 PORTLAND PLUNGER COMPARISON WITHIN-POT DATA (Q 305mm LENGTH WEAR MEDIAN MEDIAN COST/TIP COST PER TIPDAY SAMPLE MATERIAL TYPE RATIO LIFE LIFE (A$) REPAIR COST SIZE (mths) (days) NIL $75 $600 PA SPEC CAST 1.0 12 360 0.11 0.32 1.68 333 INCONEL 601 PIPE 1.1 13 396 180 | 0.45 0.64 1.53 20 HR SPEC CAST 1.4 17 504 45 KjTjH 0.23 1.21 139 NIRESIST D4 CAST 1.9 23 684 165 0.24 0.35 Q.90 90 HI CHROME (25-27%) CAST 2.4 29 864 100 0.12 0.20 0.72 49 31 OSS CAST 2.8 34 1008 120 0.12 0.19 0.62 38 INCONEL CAST 3.3 40 1188 235 0.20 0.26 0.53 30 304SS BAR 3.3 40 1188 110 0.53 22 INCOLOY 800HT BAR BKH 270 0.19 | 0.24 28 (ii) 255mm LENGTH WEAR MEDIAN MEDIAN COST/TIP SAMPLE COST PER TIPDAY MATERIAL TYPE RATIO LIFE LIFE (A$) R EPAIR CO ST SIZE (mths) (days) NIL $75 $600 PA SPEC CAST 1.6 19 576 0.07 0.20 1.06 51 304SS BAR 4.2 50 1512 80 0.05 0.10 0.42 69 31 OSS CAST 8.3 100 2988 80 0.23 100 JKS CAST 9.1 109 3276 150 0.05 0.07 0.21 26 NICROFER G3 BAR BE9I ••j^j^ 200 0.04 H*2JT^H KXES 126 WORKSHOP DATA WEAR MEDIAN MEDIAN MEAN MEDIAN STD COST PER TIPDAY SAMPLE MATERIAL TYPE RATIO LIFE LIFE (mm/m) (mm/m) (mm/m) REPAIR COST SIZE (mths) (days) $75 $600 ORIGINAL LENGTH PA SPEC CAST 1.0 11 326 5.2 4.6 4.6 0.35 1.94 2225 HR SPEC CAST 1.4 15 441 4.3 3.4 5.4 0.27 1.44 2616 HR SPEC (HEAVY) CAST 1.4 15 455 4.2 3.3 2.3 0.27 1.41 19 NIRESIST D4 CAST 1.3 14 429 3.2 3.5 1.6 0.55 1.76 16 HIGH CHROME (27%) CAST 1.5 17 500 4.0 3.0 3.9 0.35 1.38 18 HIGH CHROME(25%) CAST 2.0 22 652 6.1 2.3 9,8 0.26 1.06 11 310SS CAST 2.1 23 682 2.7 2.2 2.1 0.28 1.04 29 INCONEL CAST 1.9 21 625 3.1 2.4 3.1 0.40 1.23 19 304SS BAR 3.8 42 1250 1.3 BEIMM 1.2 0.15 0.56 20 INCOLOY 800HT BAR 2.2 24 714 2.8 2.1 2.3 0.48 1.20 17 50mm SHORT 304SS BAR 6.6 71 2143 2.2 0.7 3.5 0.07 0.31 8 310SS CAST 5.8 63 1875 1.1 0.8 1.5 0.08 0.36 317 JKS CAST 3.8 42 1250 1.1 1.2 0.5 0.18 0.59 4 NICROFER G3 BAR 15.3 167 5000 0.4 0.3 0.7 0.16 57 Notes: (0 In pot measurements were conducted every 3 months from September, 1992 to September, 1994 (ii) Workshop data is from January, 1990 to June, 1995. (iii) "Wear ratio" = (median wear rate/median wearrate o f 305mm Cast iron). (iv) Inconel pipe "metaT wear ratio was 7.0, but, due !o the Itiin pipe waits, itte "effective"fife wa s a wear ratio of 1A (v) "Cost per tip/day" = (cost per tip+tabourj/trfe in days. (vi) Bar costs assume Portland bimetal plunger design. (vii) "HR spec (heavy)" is a 20% heavier plunger with the same penetration and external dimensions as a standard HR cast plunger. The space between the top of the plunger and the cup was filled with cast iron. Cost was $50. (viii)"High Chrome" is a cast iron of 25-27% chrome. CHAPTER 9 COST REDUCTION PAGE 334 (iii) $600 repairs. As a sensitivity measure, one can consider higher costs of overhaul. In 1991, only 12% of failures for 11 integrated feeder smelters was due to pointed feeders (Table 1-V). As plants identify and design out cylinder and assembly faults (e.g. Portland), the cause of failure will become increasingly limited by plunger wear. This is typical for independent feeder systems such as D12 which had 85% of changeouts due to pointed p\unae/b. The situation is different for each plant. A figure of $600 per overhaul is used as a conservative estimate of plants which have failures mainly due to pointed feeders. Table 9-1 shows comparison costs for metal only, $75 and $600 per overhaul for 14 plunger designs trailed at Portland. Figures 9-2 and 9-3 illustrate the cost per tip/day for $75 and $600 per overhaul. Compare this to the wear ratios and life of these plungers shown in Figures 5-11 and 5-12. It is clearly evident that the plunger designs with long life become more and more cost effective as the cost of overhaul is increased. Clearly the bar materials are very attractive both in life and in costs. This is not so evident if only material costs are involved. For example, HR spec and 304SS are the same cost per tip/day at $0.09 if one considers material only, yet they are $1.21 and $0.53 respectively at $600/overhaul. Note that some materials such as Nicrofer and JKS were only tested at 255mm length. This gives a biased improved cost over 305mm plungers due to lower wear rate. Using 304SS cast as a guide, the cost/tip-day of 255mm length may be about 37% of 305mm length viz. $0.23 at 255mm length versus $0.62 at 305mm for $600/repair. However, 304SS bar was not so cost effective at $0.42 versus $0.53. This may be due to the more passive material where depth is not as significant to wear rate. Clearly, stainless plungers are more economic for plants where failures are mainly due to plunger wear. Nicrofer, Incolloy and 304SS appear to have both long life and good economic returns. Nicrofer is the preferred material (even if one adjusts for the benefit CHAPTER 9 COST REDUCTION PAGE 335 FIGURE 9 - 2 CALCULATED PLUNGER COST PER TIP-DAY AT A$75/REPAIR WITHIN-POT DATA fa) STANDARD LENGTH (305mm) PA SPEC CAST INCONEL PIPE HRSPEC CAST NIRESIST CAST HI CHROME (25-27%) CAST 31 OSS CAST INCONEL CAST 304SS BAR INCOLOY BAR fb) SHORT LENGTH (255mm) PA SPEC CAST 304SS BAR 310SS CAST JKS CAST NICROFER BAR WORKSHOP DATA (a) STANDARD LENGTH (305mm) PA SPEC CAST HRSPEC CAST HRSPEC (HEAVY) CAST NIRESIST CAST HI CHROME (27%) CAST HI CMROME(25%) CAST 310SS CAST INCONEL CAST 304SS BAR INCOLOY BAR (b) SHORT LENGTH (255mm) 304SS BAR 31 OSS CAST JKS CAST NICROFER BAR 0.00 0.20 0.40 0.60 0.80 $ PER TIP-DAY (@A$75 repair cost) CHAPTER 9 COST REDUCTION PAGE 336 FIGURE 9 - 3 CALCULATED PLUNGER COST PER TIP-DAY AT A$600/REPAIR WITHIN-POT DATA - -I: (a) STANDARD LENGTH (305mm) PA SPEC INCONEL HR SPEC NIRESIST HI CHROME (25-27%) 310SS INCONEL 304SS INCOLOY (b) SHORT LENGTH (255mm) PA SPEC CAST _: 304SS BAR 310SS CAST JKS CAST NICROFER BAR WORKSHOP DATA (a) STANDARD LENGTH (305 mm) PASPEC HR SPEC HR SPEC (HEAVY) NIRESIST HI CHROME (27%) HI CHROME(25%) 31 OSS INCONEL 304SS INCOLOY (b) SHORT LENGTH (255mm) 304SS 31 OSS JKS NICROFER 0.00 0.50 1.00 1.50 2.00 2.50 $ PER TIP-DAY (@A$600 repair cost) CHAPTER 9 COST REDUCTION PAGE 337 of Nicrofer only being tested using 205mm length). Nicrofer was only 13% of the cost/tip-day of PA spec cast iron for equivalent conditions despite having an initial cost 5 times that of PA spec. It should also be noted that most of the plungers that have the longest life also are the most economic. Thus, selection of most high nickel/chrome materials achieve all the objectives required of a plunger. Hopefully this data will convince financial personnel that the initial high cost of implementation of high quality plungers is justified. 9.4 Am SAVINGS 9.4.1 Ore leaks and Air Leaks Ore leaks are commonly due to rod seal leaks on the crustbreak cylinder of integrated feeders. The Portland Feeder Reporting System showed an 85% link between ore leaks and rod seal leaks. Plants T6, T9, T16 and T17 also have traced ore leaks to air leaks from the cylinder. Section 8.4.8 has suggested that better venting will allow the air leaks to escape and not affect air leaks; but at what cost? As noted in Section 1.2.16, air generation is a significant cost in a smelter. The cost of compressing air appears to be minor viz. about $40pa for 1 litre/min continuous flow (as calculated at Portland and verified by Ingersoll Rand). The cost of a bad (150 1/min) air leak is only $200 pa if metal can be sold, or several times that if it cannot. The average cost per overhaul of Alcoa feeders was about $718 in 1991 (Table 1-IV). The cost of 1501/min air leak is small compared to the cost of changeout, as long as the plant is not power restricted. If the plant is power restricted, the cost of leaking air is about break-even with repair costs. Changing the feeder because of air leaks is often uneconomical based on a maintenance viewpoint, however, if the air leak results in an ore leak, it is unacceptable to let the alumina fall into the pot unchecked. So there is no choice - the feeder must be removed if an ore leak results. CHAPTER 9 COST REDUCTION PAGE 338 If one accepts that it is a fact of life that rod seals will leak sometime, letting the air vent is often an economic solution to ore leaks for many plants. It is wise to use a suitable air leak test system to ensure that there is a cost effective choice of when feeders need to be changed out. This value may differ depending on the relative value of air versus power for each plant at any point in time. As mentioned in Section 8.4.8, Portland has modified each feeder top mounting flange to accommodate insulated washers (similar to that used on the mounting bolts) (Figure 8-4 and 8-5). This allows the equivalent air vent area of a 75mm O.D. vent pipe. Ore leaks almost never occur unless there is a complete blow out of the rod seal housing from the end of the cylinder. The modification cost about $40/flange and was a once- off cost. It reduced costs by $70,000 pa just on bolts/insulation usage alone. To date the only problem experienced has been alumina being pumped upwards on very bad leaks in certain situations. About 50% of flanges were modified by the end of January 1993. Installation of a 300°C rubber skirt halfway down the feeder (below the rod seal area) prevented air from going down and alumina from going up. After trials in February 1993, full installation commenced mid 1993. Examination of data for a 6 month period, showed that the frequency of ore leaks for feeders with skirts has been 4 times less than feeders without skirts for situations where a rod seal leak occurs. Thus the skirts and flange modifications effectively vent any leaking air away from the alumina. Figure 9-4 shows that ore leaks have dropped from 50/month to about 10/month. Note that there was a drop from February 1993 then a rise from June 93-January 1994. This was due to a very large increase in the number of feeders being overhauled due to installation of inserts and failures due to PA spec plungers. This had a delay effect on the perceived benefits of these changes, but one should also inspect Figure 8-8 which also illustrates the effect of the reduction in jammed objects. CHAPTER 9 COST REDUCTION PAGE 339 FIGURE 9 - 4 ORE LEAK HISTORY OPERATOR OBSERVATION HISTORY REPORT : C-CONTROL CHART Fa»dera Repaired bBTwewi 1 July 1991 and 30 SBptembw 1995 Obawvcrtion Type; 105 (ORE LEAK) 2 JUL91 0EC91 MAY92 0CT92 MAR93 AUG95 JAN94 JUN9+ N0V94 APR95 SEP95 Date Repaired Subgroup Shear n=1 Total Number eA Feeders Repair** 7355 Date Created* 2 October 1995 Total Mumb»r 6f Tesdere wMh an Observotlom 4995 Numb#r matchIng sel»erlon: 999(20%) Note; GDunts adjuated for Bam pie BJIB 9.4.2 Air Leak Test Method Venting the air in a controlled manner (so that the ore leak does not occur) is a viable alternative for integrated feeders. However, one needs a routine air leak test procedure to ensure the air leaks do not get out of control. Equipment was invented as part of this study so that Portland and other Alcoa plants could conduct routine surveys. Portland now is the benchmark across Alcoa plants on the lowest air leak rate at 6 1/min per feeder; a very low figure compared to most smelters. These checks are now being done monthly to ensure that the air leak control is an ongoing activity. The method to test for pot air leaks is to shut off air prior to the DCV via a ball valve. Measure the pressure immediately, then again a period of time later - say 30s. The rate of fall of pressure will indicate the air leak rate by use of the gas law viz. CHAPTER 9 COST REDUCTION PAGE 340 PiV, = P2V2. Vj is the volume of the pipes between the valve and the piston, PI is the initial absolute pressure when the ball valve is closed and P2 is the final absolute pressure after a period of time t. The change in volume over time t is the leak rate. This technique is easy to implement, costs very little (about $60/tester) and can assist greatly in minimizing air leaks in the potrooms. D10 had a campaign to reduce plant air usage in 1992. They found the largest air leaks came from their 200mm O.D. crustbreaker cylinders. They changed out their worst ones and dropped the usage by one compressor; a saving of $200,000 pa. D12 conducted a similar investigation in 1993 and also reduced air consumption by 75,000 1/min. The main cause was faulty DCVs. Now D12 plant air pressure is much higher than before. Implementation of routine pot air leak testing at Portland has identified the main causes were by-passing DCVs, leaks on air hoses and incorrect setting of feeder isolation valves. Crustbreaker cylinder leakage was not a major cause. 9.4.3 Boyne Feeder Maintenance Strategy Boyne use a preventive maintenance strategy for their feeders based on air usage (Dunstan 69). Their cylinders (Parker) leak very badly. Average air leak rate has been found to follow the empirical equation: y = 0.37e017x Where "y" is the leak rate in m3/hr and "x" is the age of the point feeder in months. CHAPTER 9 COST REDUCTION PAGE 341 Feeders were changed out when air leak rate was such that there was insufficient air available from the plant air compressors. On this basis, all feeders in a pot are replaced after 22 months based solely on air usage. This relates to a leak rate of 70 1/min per feeder (cf. 90 1/min for a Portland feeder to create an ore leak and average Portland air leak rate of 61/min). Boyne believed there were several attractions to this procedure: (i) controls air leakage to a manageable level, (ii) allows scheduling of feeder replacement weeks in advance. (iii) the location of each feeder and its age is known. (iv) the work can easily be contracted out. (v) changeout procedures for improvements can be scheduled. (vi) scheduled overhauls have led to significant price reductions for components. (vi) ensures uniform feeder age. Problems with this philosophy are the lack of knowledge of the life of parts (as none fail) and, "just to be safe", all parts need to be changed out. Overhaul costs are higher than necessary. It is preferable to identify air leaks on pot-to-pot checks on a routine basis and, unlike Boyne, identify which of the feeders in the pot is faulty and repair only that unit. Do not change all for the sake of one. The other units may have years of life left. Boyne's procedure is not recommended as it is based on an acceptance that severe air leaks are inevitable. It is interesting to note that in 1993, Boyne started trials on retrofitting Terry seals that were a design based on the author's recommendations for Portland. These seals were successful to such an extent that by mid 1994, a decision was made to retrofit Parker cylinders with Terry seals. Once again, another plant has found Teflon seals are preferable to Viton. Also, the same basic seals are now used in Portland, Boyne, TIO and trial units at Dl 1. CHAPTER 9 COST REDUCTION PAGE 342 9.4.4 Non-routine Cylinder Overhaul Strategy Analysis of air leaks on cylinders as they enter the workshop can be used as an aid to a needs-basis overhaul strategy. The overhaul philosophy to only service cylinders with an unacceptable leak, has several risks: (i) What air leak rate at room temperature is unacceptable at operating temperature in the pot? (ii) How fast will the air leak rise in service? (iii) Will the temperature cycle of the cylinder result in premature leaks in the plant? (iv) Will the tradespersons feel that management is accepting poor quality out of the workshop and, in future, not be accountable for their workmanship? The last item is the one that is the most serious question, as poor attitude by tradespersons can quickly create a major quality problem with feeders. At Portland there has been a major improvement in the dedication and quality control of tradespersons working on feeders; to lose this positive attitude would be most undesirable. Feedback of the results of trials has shown tradespersons the success of this procedure, so there has not been a problem despite initial misgivings. This is an example of the benefits of an open attitude to discussion between tradespersons, operators, engineers and supervisors. A trial was conducted at Portland to see if air leaks were different at pot temperature and at room temperature. Ten feeders had their air leak rate measured on the pot. They were then removed and tested again at room temperature. This showed six of the ten had higher air leaks at room temperature, none had less. During the cylinder trial, seal leak rates increased 7 times at room temperature when cylinders had cooled down (Section 7.4.3:10). CHAPTER 9 COST REDUCTION PAGE 343 Hence, repairs of cylinders due to any air leak at room temperature may be inflating the cost of overhaul. In the example above, all 10 cylinders would have been overhauled at a cost of $600 each, despite the fact that the feeders that were removed were operating quite satisfactorily before the test. The average Portland feeder leak rate was 6 1/min, or $ per annum, yet Portland used to repair 85% of feeders due to "excessive air leaks" at about $600 each. A procedure and equipment was designed by the author for the workshop using the previously mentioned Gas Law method (Section 9.4.2). Gauge capability tests showed that the repeatability and reproducibility for the method was statistically capable of a tolerance region of 0-30 1/min. The cost of the equipment was less than $200, but the assistance it gave to cost effective maintenance was many times greater.. Figure 9-5 shows a flow sheet comparing costs and life of feeders with "full" (cylinder) overhauls and "quick" (no cylinder) overhauls. Full overhauls only achieved a 10% greater life but they did so at a cost 3.6 times that of quick overhauls. Also, the cost of feeders on the second overhaul was not different to the previous quick/full overhaul. Since the installation of testing equipment in the workshop in January 1993, there has been a 30% reduction in costs and little change in the life of feeders. Clearly, this strategy has major cost attractions. This is a great opportunity for cost savings if proper air leak management and adequate venting is provided. Three-monthly air leak checks on the pots showed that there was no problem with feeders having excessive air leaks in the pots despite the cylinders not being overhauled if the leaks in the workshop were under 30 1/min. The concern of escalating air usage in the plant proved not to be valid, and the tradespersons concerns were put to rest. This strategy of overhaul is recommended. CHAPTER 9 COST REDUCTION PAGE 344 FIGURE 9 - 5 CYLINDER OVERHAUL STRATEGY COSTS AVERAGE QUICK AVERAGE FULL OVERALL COST = $189 ea OVERALL COST = $688 ea QUICK OVERHAULS Vs FULL OVERHAUL * 10% LOWER LIFE * 70% LOWER COST NOTE: Portland feeders with Terry cylinders overhauled 1/6/93 to 31/7/94 CHAPTER 9 COST REDUCTION PAGE 345 9.5 SAVINGS FROM REUSED FEEDERS If the feeders generally work satisfactorily the day the pot is taken off-line, then it is not necessary to fully overhaul them. The hazards are almost the same for changing feeders in a live pot as in a dead pot. The only difference is temperature as dust, fumes and potential falls are about the same. For pots that have been taken off-line due to excessive temperature, there is a risk that the feeders may have been overheated. It is wise to overhaul these feeders. These pots are the exception to normal pot failures (Section 8.4). The practice of not overhauling feeders from off-line pots should be considered for the following reasons: (i) At Portland, "dead pot feeders" in 1991 cost about $420,000 to overhaul. Alcoa spent $1 million. Probably at least half of this could be saved i.e. they will be "average" feeders. (ii) Protect new feeders from temperature spikes - so when they are installed in a pot they stay longer (refer Section 8.4.2:4). High temperature can rapidly reduce cylinder life. (iii) Higher temperature of new pots tends to tighten up seals and bushes which slow feeders (refer Section 4.5.2) This causes dagging of feeders on new pots, blocked feeders and unnecessary exposure of operators to dust, heat and fume. (iv) New integrated feeders are slower and have increased shot size variability (refer Section 2.5.4:3). This is aggravated by expansion of components at high temperature. (v) Smaller plungers (on old feeders) give greater pressure at the crust. New pots have harder crust to normal pots due to lower calcium fluoride content and higher calcium fluoride/aluminium fluoride ratio. Greater pressure will assist in preventing blockages. (vi) Smaller plungers tend to stop plungers jamming between the large anodes that are in new pots. CHAPTER 9 COST REDUCTION PAGE 346 Portland initiated reusing old feeders in new pots for 3 months in 1993 and saw no change in feeder life. From January 1994, the practice was started again (after installation of inserts in all pots). By mid 1994, there was a savings of 10% in costs and a reduction in feeder changeouts by 10% that could be directly attributed to this practice. This practice has been used by Tl 1, T16 and T17 with similar success. 9.6 FEEDER TRACKING SYSTEMS As there are many pot feeders in a smelter, it is difficult to determine the link between overhauls and failure. A major benefit of a good pot feeder tracking system is the identification of cost reduction opportunities and the ability to follow a trial through to completion. Often the person who initiates a trial does not know the results, so trials are often repeated or lapse. However, if there are no failures, one does not need a tracking system, therefore, many independent feeder plants may not need a tracking system. 9.6.1 The Purpose of a Feeder Tracking System At Portland operators changeout feeders, but, in most plants, tradespersons do all changes in the potrooms. Overhauls are usually conducted in a workshop remote from the potrooms by tradespersons. Hence, there may be up to three groups (or "links in the chain") which may have little to no personal contact with each other. Each group are often most ignorant of how the feeder is being used and his/her effect on how it operates. The supplier and user seldom see (or even know personally) each other as they are working in quite remote locations The "reason for removal" may appear to be totally unrelated to the "mechanical fault" observed by the repairer as the repairer does not have the feeder at elevated temperatures and there will be no alumina present as this is usually cleaned out before it reaches the workshop to keep workshop housekeeping to a high standard. Once overhauled, the feeders are very seldom checked (even at room temperature) and are almost never checked with alumina to see if they dose correctly. Although this may CHAPTER 9 COST REDUCTION p seem a pedantic discussion, it is very important when discussing feeder failures because the tradesperson is not aware of operational situations, and this remote sequence is one of the flaws in effective feeder performance. An example of this is the experience at Portland up to mid 1990 where about 30-50 feeders per month did not have any repairs conducted by the tradespersons when they arrived at the workshop; nothing could be found wrong with them except for signs of mechanical damage or foreign objects were jammed in the dosing unit; often superstructure mounting bolts or washers. The operator was frustrated in what was perceived to be shoddy work because the feeder internals were vibrating uncontrollably. The repairer considered that the operator was incompetent as there had been no repair to the feeder and it seldom came back a second time. Once a tracking system was installed where the reason for removal was linked to mechanical faults observed in the workshop, it was discovered the cause was incorrect installation of feeders in the superstructure. If alumina was not swept out of therim of the superstructure (Figure 1-11), the feeder sat on a bed of alumina. The alumina gradually vibrated out of the rim area after a period of operation so the mounting bolts became loose. This in turn resulted in the bolts falling into the alumina and consequently jamming the dosing spool. The operators generally correctly tightened the mounting bolts, but only the most experienced ones were aware of the "trick" to clean out therim. Bot h parties were completely unaware of that their effect on each other. Once the link was made, training and reinforcing of the mounting bolt washers dropped failures from 35 to 25/month (Figure 8-8 in 1989-1990). Despite the frustration by both parties, other benefits were a reduction of $3,000/month in repair and transport costs. It is very important that a tracking system is used to trace feeder performance in order to link all the people involved otherwise one may be solving the wrong problem. Up to 1990, most feeder development was centred on finding a better plunger material despite CHAPTER 9 COST REDUCTION PAGE 348 the fact that 80% of failures were due to ore leaks. The engineers were not aware of why feeders were removed. They were only aware of what wore out by looking at feeders in the workshop, so were working on the wrong things. They should have been working on air cylinders until they were under control, then work on plungers when the causes of ore leaks were identified and fixed. Thus it is important to ensure that any "reason for removal" code does not have any words that are ambiguous and does not imply the cause of failure. For example, one can observe that the feeder is leaking, but one cannot "observe" that the spring is faulty as one cannot see the spring. For the tracking system for Portland the reason for removal is called an "operator observation" and does not assume what are "mechanical faults" which are independently reported by the repair tradespersons. This terminology is recommended. 9.6.2 Portland Feeder Reporting System Since 1987, each feeder overhauled at Portland has had the cylinder numbered and the parts used listed against the cylinder number. The system was significantly upgraded in 1990 to include feeder assembly numbers, plunger numbers and location in the pots. With the inclusion of statistical analysis, the Feeder Reporting System (FRS) is possibly the most comprehensive feeder tracking system in the world. T8 and T17 have reasonable systems, but not as elaborate as an engineering, analytical, accounting and inventory tool which has operator and tradespersons directly interacting with on-line screens. Following the author's Feeder Recommendations to Alcoa plants in 1992, all Alcoa plants have introduced some type of tracking systems, but none as complete as Portland. Note that a plant with low turnover of feeders probably does not require a tracking system as there is little short term benefit. However, any plant with a less than acceptable maintenance budget or one that wishes to optimize feeders should seriously consider a system similar to that at Portland. CHAPTER 9 COST REDUCTION PAGE 349 Portland has established a computer reporting system on the process VAX computer system. Computer screens are located in the workshop (for tradespersons) and in the potrooms (for operators). In this way, both groups directly affected by the feeder have an input into its performance. The FRS incorporates the following main items with the ability to add comments if needed at every step in the process: (i) Feeder LD. Each cylinder and feeder assembly are uniquely identified by a distinct code number for tracking performance of parts, costs and modes of failure. Plungers are coded with the date of casting and the batch number of that day. Each batch has a chemical analysis from the supplier. (ii) Test after overhaul After overhaul, each cylinder is tested for cushioning (by use of an accelerometer), seal leakage (flow meter), plunger O.D. (calipers) and a number of other minor items. Comments can be logged for the test data. (iii) Installation On installation in a pot, the operator enters the cylinder number, pot, location, date and crew. Hence, each feeder is uniquely identified in the potrooms. (iv) Removal On removal, the operator identifies the cause of removal (called "operator observations") against the cylinder number. (v) Overhaul When the tradesperson receives a feeder to overhaul, the cylinder number is entered onto the computer. The tradesperson is automatically shown any operator observations, any comments, the age of critical parts and a list of recommended parts needed to be changed irrespective of the reason the feeder was removed from the pot. This information may be based on Weibull analyses of past data. In this way, the people who service the equipment are aware of the reason for overhaul and any items that need changing irrespective of failure. This avoids early failures of items that have almost failed e.g. if partly fatigued. CHAPTER 9 COST REDUCTION PAGE 350 At overhaul, the tradesperson carries out about 20 standard checks of the feeder unit and lists "mechanical faults". The feeder is then overhauled and all repair items are entered along with whether they are "failures" or "preventive maintenance". (Failure and preventative maintenance data is used for Weibull analyses.) Plunger O.D. is measured at overhaul. If the plunger is of acceptable size, it is reused. In addition, if a feeder fails within 100 days, an alarm is raised to highlight that extra attention is needed on this feeder. From this database, a complete history is collected for each feeder (or component) of all parts and where it was placed in the pot. Enquiries are possible by component, by pot, or by groups of pots to see when services have been done, the operator observations, mechanical faults, comments, plunger size, plunger wear rate, cost of overhaul, parts used, current part life, dates of test/install/remove/overhaul and operator/tradesperson who completed these actions. Reports of feeder performance, cost control and part life are available in tables and statistical output packages viz. Pareto charts, control charts, histograms, bar charts, Weibull analyses. This allows detailed assessment of cost and performance control by engineers, accountants and overhaul contractors. The database started in 1987, but system integrity improved significantly from June 1991 when the PC data were transferred to the VAX system and standard reporting procedures were satisfactorily established. Trades entry of parts was commissioned in August 1991 and the operators in September 1991. 9.6.3 Tracking Systems and Quality Control Control charts, Pareto charts, Weibull analyses, tables of costs and tables of turnover rates of parts enables examination of high frequency and/or high cost items in a non- biased manner. Use of time related graphs such as control charts indicate if changes have achieved the desired results. There are several examples in this document viz. CHAPTER 9 COST REDUCTION PAGE 351 Figures 4-6, 5-3, 5-10, 7-5, 8-3, 8-7, 8-8, 8-9, 9-4, 10-3, 11-2 (top) and 11-3. A great deal of the data presented in tables has been generated from the tracking system. Some examples of problems identified by tracking systems and quality control checks include: (i) After quality checks ex the workshop were initiated, new Terry piston seals at Portland were found to be 50% of the design thickness and had been for 5 years. Checks at T6 found the same problem for the previous 20 years using Terry cylinders. (ii) The clearances of Terry bushes and seals at Portland were too tight and affected feeder speed. A similar problem was seen at T16. (iii) Spigotted flanges installed to improve alignment at Portland proved to have no effect despite "good engineering judgment". Misalignment in the spool had a bigger effect which was unnoticed until traced using the tracking system. (iv) Backup washers on Parker pistons cost more than the seals they were installed to protect. (v) Previous focus to reduce seal costs was changed to optimizing overall repair costs. Seals only cost about $10 but can be responsible for repair cost over $700. It was more efficient to pay more for seals to prevent downstream costs. (vi) Plunger O.D. measurements identified the wear rates of plungers and led to better plunger materials. (vii) Failure rates were linked to pot conditions. The major cause was the pot, not the repair quality. (viii) Weibull and 10/50/90% failure analyses gave early indication of the effect of design changes rather than having to wait for total failures before making decisions on the success of trials. (ix) Tradespersons and operators became more aware of costs and the impact of their jobs resulting in better quality products resulted. CHAPTER 9 COST REDUCTION PAGE 352 These are only a few of very many benefits obtained by use of feeder tracking systems. By far the greatest benefit is the ability to identify the effect of changes. If a modification is installed, one can compare the benefits against a "control" sample sent at the same time from the workshop. For this reason, all trials that are conducted at Portland have a "control" group as well. To ensure there is an adequate sample size for statistical comparisons of trial versus control, any trials are conducted on 20 or more feeders. Thus, the results of trials are quickly either (i) accelerated to production changes or (ii) the modification is stopped. In all cases, data are used to establish the findings rather than use subjective judgment. CHAPTER 9 COST REDUCTION PAGE 353 9.7 MAJOR FINDINGS FROM COST REDUCTION (i) High nickel alloys are more cost effective than less passive materials despite the higher initial cost. (ii) Bimetal and shorter plunger designs optimize the material cost of the plungers. (iii) Piston and rod seal air leaks contribute to ore leaks (for integrated feeders) and to costly air consumption (integrated and independent feeders). (iv) Non-routine cylinder overhaul is a cost effective alternative to 100% changeout of seals and plungers with little change in mean time between failures. (v) Overhauling feeders from off-line pots is not cost effective. (vi) Tracking systems are powerful tools in optimizing designs and costs. 9.8 RECOMMENDATIONS FROM COST REDUCTION (i) Install cast JKS or Nicrofer plungers using the Portland bimetal design. (ii) Trial low penetration plungers or drop bath levels so plungers do not get wet. (iii) Check air leaks often and provide air leak testers for troubleshooting. (iv) Overhaul cylinders based on quality control guidelines that include cylinder bypass and rod seal air leak rates. (v) Do not overhaul feeder from off-line pots. (vi) Install a feeder tracking system that links where the feeder was located, reason for removal, mechanical faults and parts changed. CHAPTER 10 FEEDER DEVELOPMENT PAGE 354 CHAPTER 10 FEEDER DEVELOPMENT 10.1 SUMMARY In parallel with optimising the existing feeders at Portland, this research investigated integrated and independent feeder designs. Three feeder designs were invented by the author and two have been successfully patented. The A2 integrated feeder proved to be unsuccessful. The Pulse Chute did not progress past the prototype stage due to interest in the A3 independent feeder design. The A3 feeder incorporates the advantages of the independent design but can be retrofitted on line to many potlines that have integrated feeders at about half the cost of retrofitting traditional independent feeder designs. A cheaper Sequential Feed design has been developed that is a compromise between the A3 and traditional Alcoa AEDD feeder designs. The new feeder designs incorporate many features that enhance long life and ease of maintenance. To date, operating results on the trial pots are encouraging. Comparison of integrated and independent feeders using data and observations from this research shows that the independent feeders are superior in almost all parameters except capital cost. CHAPTER 10 FEEDER DEVELOPMENT PAGE 355 10.2 PROJECT HISTORY From early 1990, development commenced on feeder designs that achieved long life, direct feed and reduced maintenance costs. By mid 1990, the A2 feeder was invented (Appendix 6) with trials in the plant commencing in February 1991. Patents have been approved in Australia, New Zealand and the USA. Improvements to the shot size accuracy of the existing AEDD feeder (by development of the spool insert shape) was a byproduct from this development. Late in 1991, the Pulse Chute was developed and reached prototype standard by late 1992 (Appendix 7). As the A3 has significant benefits over the Pulse Chute, the chute did not proceed past the prototype stage. The A3 was invented late in 1992 (Appendix 8) and plant trials started late in 1993. The modifications in mid 1993 to the assembly of the existing AEDD to stop jamming, was a direct result of the A3 development. Figure 1-22 shows all these designs and illustrates that the overall dimensions are similar so they can be installed in pots where AEDDs are currently installed. This was one of the original objectives. Table 10-1 shows a comparison of the different feeder designs against features that are important for feeder or pot operation. Table 10-11 shows comparison plant scale data of the trial feeder designs compared to AEDD pots over the same test period. From mid October 1993, trials commenced on seven pots at Portland. This was a "disaster check" to see if these designs had any major drawbacks. Realising the variability in pot operation, a trial of one or two pots of different design is never sufficient to prove or disprove a feeder design or pot operating strategy; a trial of about 10 pots would be required. However, a small trial would indicate if the results were very good or very bad and would assist in tuning equipment and electrical components and to debug software. Based on this small trial, it was envisaged that a larger trial would follow when funds were available. Below is an explanation of the designs, costs and benefits (to date) of feeder development at Portland. 1- Z LU Q _^ co Z o- co CO CM CO CO CM m < LU 4^ d UJ T— CM d d 0- c^- d UQJ- m Q Z _J Q < UJ CO CO CM in d CM & LU - c^- o d d • d AED D o >- CO QUENT I UJ 1LU- d CO Z H >- Z UJ LU Q CO CO z CO co CO in CM CD LU i UJ "*" 1 T— Z d CM CM o m o QL >- UJ LU d CJ) 0_ Q Z a LU HH - m m csi CM co o CM co < H § • a. (trofi t c £ x: a. o <1- o >cu o co 3 MC—O cu !t 1— TJ TJ sO C o o CU K % CU CU TCJU > cu 2% CO s 4—I ro CQ I— 0 ro TJ TJ cu CD cu O -Q > CU cro 4-rof _c «E> cu < Q UJ UJ h- _J 1- LL >- O r- LU LU LU r- O o 2 co < < t CD z o UJ z I— CaO. Q CO C O (OF F UJ ;* STAN I LOU 2 z LU 3 < LUL LU CO LL < = o oc CO < LU DC LU IO N LL. ;IZ E LU z tr H o_ o > Z)\- z< Uzl r- z CO CO CO H UJ LU TABLE 10-H FEEDER TRIAL DATA PARAMETER A2 SPLIT HEADER SEQUENTIAL FEED A3 (1075,1085,2069) (1022,1094) (2075) (2070,2071) JULY-AUGUST 1994 ALUMINA IN POT - PERCENTAGE LOW - STANDARD DEVIATION GOOD ANODE EFFECTS PER DAY OK OK GOOD FAIR ANODE EFFECT DURATION TIME OK OK OK OK PRODUCTION RATE - mm TAPPED PER DAY OK OK FAIR OK -TONNES PER DAY OK OK OK OK APRIL-JUNE 1995 ALUMINA IN POT - PERCENTAGE LOW LOW - STANDARD DEVIATION GOOD GOOD ANODE EFFECTS PER DAY GOOD FAIR ANODE EFFECT DURATION TIME FAIR GOOD PRODUCTION RATE - mm TAPPED PER DAY OK GOOD -TONNES PER DAY OK GOOD Notes: (i) "GOOD/OK/FAIR" rating based on where pot rated against a range of 28 similar pots. "GOOD" means the pots were in the top 25% of 60 pots in the first period and 28 pots in the second period. Similarly, "FAIR" were in the worst 25% and "OK" were in the middle range. (ii) The lower the concentration of alumina in a pot the better the current efficiency but the harder it is to control operation. "LOW concentration would generally indicate better production rate but potentailly more tendency for anode effects. (ii) Between the first and second comparison period, the plungers on the A3 feeders were increased in length to increase penetration which may be a cause of previously average performance on anode effects. (iii) Production rate is expressed by two independent methods, "mm TAPPED PER DAY" is the distance the anodes moved during tapping. There is a linear relationship between mm tapped and production rate. This the most accurate method of assessing production rate. "TONNES PER DAY" is measured by the crane that hold the crucible. This parameter may be affected by any bath that is mistakenly tapped. Over a long period and by using comparison data for a group of pots, these data can be used for production rate. (iv) In the second period, there are no data for A2 pots as all were changed back to AEDD design. The split header pots were not considered dueto thei r average performance in the first period. (v) Dueto th e small sample size of trial pots, these results can only be used as an indication of trends. Excursions on such a small sample size can greatly affect comparison results. A much larger sample size for an extended period is requiredto compare feeder designs. CHAPTER 10 FEEDER DEVELOPMENT PAGE 358 10.3 A2 FEEDERS, DRIBBLE FEED AND PULSE CHUTE The A2 feeder is an integrated feeder that looks like the AEDD, but feeds when the plunger is up, not down (Figure 1-14). From late 1990, up to 8 pots have been operating with A2 feeders. Performance has been quite variable from pot to pot (until mid 1993) due to fouling of dags on the single feeder chute (shown in Figure 2-1). In mid 1993, the chutes were changed to thefinal double outlet design (Figure 2-17) and performance improved with results comparable to normal pots. Overall, the benefits of direct feed, accurate shot size and smaller shot size have not been evidenced possibly due to tuning factors in the feed control logic. It is also possible that any improvements to performance are not sufficiently large enough to be seen. Thus, the cost of retrofit is not justifiable. The main operating problem with the A2 feeder is that any dagging can foul the chute, then the pot is starved of alumina. It was decided late in 1993 that this design was not advanced enough to continue trials. These pots are now being converted back to AEDD as the pots come off-line. Trials were conducted on a dribble feed continuous feeder chute in 1990-1991. This proved unsuccessful due to blockages of the small orifices from tramp material. This material either came from the alumina or was drawn up the nozzles from the pot via dags. Problems with blockages are common in continuous feeders and were a major reason that other companies also dropped continuous feed. Comalco continues trials on a pot with continuous feed at Bell Bay. To date the trials have not progressed past a one pot trial despite the opportunity presented by the current expansions at Boyne Island and NZAS. Given that continuous flow is prone to blockages, the alternative is to pulse the alumina in short shots of a larger quantity. The Pulse Chute distributed a metered dose of alumina on an intermittent basis over time (Figure 1-18). In this way, a semi- continuous feed can be achieved, yet the problems of blockages were not a factor. This CHAPTER 10 FEEDER DEVELOPMENT PAGE 359 could be achieved by an airslide with 5-10s pulses of aeration. As it would no longer be necessary to have very small orifices to control flow, large outlet nozzles could be used to distribute the alumina thus avoiding blockages. The design was taken to prototype stage but then dropped in late 1992 due to interest in the A3 which was considered a better alternative. The pulse chute may also have problems with alumina sifting into the air plenum when there is no pressure in the plenum. Although purge holes are part of the design, this is still of concern. Although patents were applied for world-wide for this design, these applications were withdrawn in early 1994 due to the high cost of renewing applications for a design that was unlikely to be taken into production. However, the design of the chute outlets assisted in the development of the double outlet chutes used for A2 and A3 feeders (Figure 2- 17). 10.4 MULTIPLE DCVS PER POT OPTIONS One major problem with the existing Portland AEDD feeders is that five feeders operate from the one valve. This results in: (i) lack of air when all feeders operate, leading to - poor shot size accuracy, - different feeder speeds and - dwell time time-out before the slowest feeder releases dose and/or breaks a hole, (ii) time delay across the pot for feeders to operate, resulting in long dwell times and subsequent dags, (iii) high air consumption as much air is used tofill the air line eg. 40% at Portland. CHAPTER 10 FEEDER DEVELOPMENT PAGE 360 10.4.1 Split Header If the existing air header is split in two, then two valves could be used to minimise the effect of the above factors. This design is called a split header (Figure 10-l(a)). The two valves operate at different times. Rather than a feed every 3 minutes, a feed could occur in half that time interval somewhere in the pot. If any mixing occurs across the pot, then the pot alumina concentration would be more consistent over time. There is usually a great deal of mixing across a pot due to the agitation of the bath and metal. Two designs are under trial. One has both valves on the wall of the building and one has the valve on the pot itself (with the exhaust piped into the pot gas cavity) to minimise air usage. Results to date from the two pot trial suggest there is not a significant improvement to pot performance for this design. 10.4.2 Sequential Feed Sequential feed design is the provision of a separate DCV for each of thefive feeder s so that each feeder can operate at a different time across the pot (Figure 10-1(b)). This gives more consistent alumina concentration in the pot over time, allows more repeatable feeder speed, less dwell time and more accurate shot size. As with the A3 feeder, the only changes necessary to the pot are installation of electrics across the top of the pot and modifications to the PLC program. The impact on production is zero for a retrofit of this design. Shot size accuracy is similar to the A3 feeder (Section 2.4.5). This design has many of the benefits of the A3 but at a cheaper cost. There is only one pot currently on trial. Results show good anode effects per day and good control of alumina in the bath. A larger sample size is needed to gauge pot performance. The hardware has been faultless up to date viz. 20 months. CHAPTER 10 FEEDER DEVELOPMENT PAGE 361 FIGURE 10 -1 MULTIPLE DCVs PER POT FEED OPTIONS A. SPLIT HEADER 2 DCVs/POT a existing feeders „ 2 or 3 feeders operate at one time existing air header split in two by new valve EXISTING DCV NEW -51 \ l • ,DCV PLUG 000QQ EXISTING WALLBOX B. SEQUENTIAL FEED 5 DCVs/POT a existing feeders „ one feeder operates at one time DCV WALLBO"0000X 0 -- FIGURE 10 - 2 INDEPENDENT FEED OPTION FEATURES • five crust break cylinders operate at different times • crustbreak and feeding occur at different times • new five feeding cylinders (with hole) under existing cylinder • feed cylinders operate sequentially across the pot eg. 1,3,5,2,4 • each cylinder can be turned off if required eg. after set, hole mucky J3 FEED g DCV O O O O Q FEED CYLINDER PLUG CRUSTBREAK WALLBOX DCV CRUST BREAK O00O© CYLINDER CHAPTER 10 FEEDER DEVELOPMENT PAGE 362 10.5 A3 FEEDER The A3 independent design is believed to achieve most (if not all) the original objectives of this research project (Table 1-m and Section 1.3.1). Shot size tests showed excellent repeatability and control (Section 2.4.5 and 2.6.4). Pot installation included the sequential operation of feeders across the pot so that not all are fed at the same time (as is the case with Sequential Feed) and there was a delay between crustbreak and feed (Figure 10-2) so that the alumina flowed into the open hole. This section will discuss the results of the A3 trials to date and its design features. 10.5.1 A3 Pot Performance The A3 was invented, scoped and commissioned by the author. Matthew Langmaid and Colm Fitzpatrick (Mechengineering Services) designed the feeder on CAD and precommissioned the prototypes. All three people worked as a team to develop the final design. The author appreciates the skill and cooperation of Matthew and Colm in achieving what has proved to be a well designed piece of equipment. Although there have been several problems with some of the dosing cylinders seizing due to too tight seals/bushes (but never on Parker cylinders) to date (1.11.95), no mechanical faults have occurred to any part of the A3 design itself or on any crustbreak cylinders on any of the 10 feeders in almost 2 years of operation. This is most unusual for Portland feeders as 20% were failing in 3 months and 50% failed in about 3-4 years when installed at the same time. One would have expected 2-5 to have failed by now. It is even more unusual considering the extra equipment added to the feeder over the standard AEDD. Two pots were changed over from AEDD to A3 feeders in December 1993. Results to date (Table 10-11) indicate the A3 pots has very good control of alumina concentration and standard deviation of concentration in the bath. Median metal production is in the top 25% of pots, but the anode effect rate is higher than normal. This may be because the concentration is running close to anode effect all the time. This is desirable for metal production (which is evidenced by the data) but tends to CHAPTER 10 FEEDER DEVELOPMENT PAGE 363 increase anode effects per day. It may also be due to some sticking of the dosing cylinder for some designs. There has been some problems for some cylinder designs. Operators often cannot see any problems when the pot has anode effects. Hence, it could be lean operation or cylinders sticking (or both). Further investigations are needed. It is pleasing to note that, even though the frequency of anode effects may not be in the top 25% of pots, the time in anode effect is small. Once again, if the heat balance was poor, then the metal production rates would not be in the top 25% of pots. As time in anode effect was small, the generation of potentially global warming gases would be less than normal pots. As was mentioned with the split header and sequential feed trials, the sample size is too small to be specific on the ratings of these trials. One needs a large sample size to make any definite conclusions. No problems have occurred with the feeder design itself. To date, it appears most of the predictions on the A3 design have proved correct with the small sample size available. The main eauipona-tr- problems have been with seizing of the dosing air cylinder at elevated temperatures and some valve failures (Section 10.7). Cylinder seizing was due to insufficient clearances on the piston seals on some brands of trial cylinders. There is still a concern that this sticking problem may still exist intermittently and may be the cause of higher than normal anode effect frequency. 10.5.2 Possible Use in Other Smelters The basic design of the A3 is to cut a section out of the existing assembly and replace it with a feed cylinder that has a hole in the centre of the piston rod - a "hollow cylinder" (Figure 10-3). The cylinder operates the dosing unit instead of the spring/collar operation of the AEDD. Figure 1-15 illustrates how it operates and Figure 10-3 shows how the feeder looks compared to an AEDD feeder. CHAPTER 10 FEEDER DEVELOPMENT PAGE 364 FIGURE 10 - 3 A3 AND SEQUENTIAL FEEDERS • CYLINDER CRUSTBREAK CYLINDER^ FEEDER DOSING _ ASSEMBLY CYLINDER DOSING DOSING AREA UNIT PLUNGER PLUNGE & SHAFT A3 A3 AEDD ASSEMBLED DISASSEMBLED AEDD AEDD A3 SEQUENTIAL FEED CHAPTER 10 FEEDER DEVELOPMENT PAGE 365 The A3 design was developed to incorporate the benefits of the independent feeders, but utilise Alcoa integrated feeder equipment. As such, the design can be retrofitted in 15 plants using Alcoa feeders - Portland, T2, T3, T4, T5, T6, T7, T9, TIO, Til, T13, T14, T15, T16, and T17. It could be installed in T8 with some modifications. Cylinder dimensions have been chosen such that the one design of cylinder will fit in all plants, thus reducing the effective cost for mass production, allowing common spares between plants and ensuring the results of trials conducted at Portland are directly transferable to other plants for quick implementation. Estimates to install A3 independent feeders in pots at Portland (at 5 feeders per pot) suggest that the capital cost would be about half that of retrofitting conventional independent feeders, mainly because the cost of superstructure is much lower on a retrofit and significantly lower on a new pot. At this time it is not known if the cost is justifiable, but current trials at Portland are afirst step to quantify the cost benefits of independent feeders over integrated feeders. If there is a cost benefit, then the A3 is a cost effective alternative in order to retrofit or to install in new pot. In addition, the A3 can be retrofitted on-line, whereas the integrated feeder needs to have the pot off-line due to significant superstructure modifications. 10.5.3 A3 Design Features The A3 has several attractive design features that have been incorporated from experience of the AEDD development and A2 performance. (i) Spigotted Flanges The assembly design incorporates spigotted flanges at every join and bronze bushes are used on moving parts to minimise misalignment. Normal Portland AEDD deflection of the plunger is 25mm side to side; the A3 feeder achieves 5mm which is comparable to other independent feeders (Table 8-III). CHAPTER 10 FEEDER DEVELOPMENT PAGE 366 (ii) Shot Size Accuracy Shot size accuracy has been proved to have a standard deviation 3-4-Ws better than the existing feeders (Section 2.4.5 and 2.6.4). These results may be similar to that achieved by other independent feeders as the cylinder sizes are similar. However, the mass flow design of the A3 should give superior results. Although not mandatory, the trial A3 pots have double outlet chutes rather than the normal single outlet Alcoa design. This is to give more direct entry of alumina into the hole. This may give better results than the AEDD dust loss (27) and achieve benefits claimed by Reverdy (26) that direct feed is good for dissolution (Section 1.2.5). (iii) Maintenance The design was developed with maintenance in mind. It is possible to remove any section of the unit with minimal effect on remaining parts. For example, it is possible to service the dosing unit without disassembling the crustbreak or dosing cylinder. It is possible to do air leak and piston bypass checks without major removal offittings on the top flange. (iv) Ease of Installation The top flange equipment (including the valves) is no higher than the top of the superstructure to avoid any possibility of being hit by an overhead crane. No modifications are necessary to the superstructure other than to run conduit and wiring for the electrics to the valves. No air line changes, oxy-cutting or welding are required. These features make it easy to retrofit in operating pots. (v) Cost Control It is cheaper to install the A3 feeder than the other independent designs, as there is less superstructure work required. The costs of valves, cylinders, electrics, programming and feeder assemblies are probably similar for the two design types, but changeout and maintenance costs are less for the A3 as it is more compact and easier to service. CHAPTER 10 FEEDER DEVELOPMENT PAGE 367 In summary, the A3 feeder achieves at a lower cost the features of independent feeders viz. low air usage, direct feed, feed and break at different times, immediate crustbreak pressure, low dwell time, sequential feed across the pot, direct control of dosing unit, accurate shot size. In addition, the A3 feeder can be retrofitted in existing smelters with little to no effect on operations. Thus it is a cost effective competent design. 10.6 END STROKE SENSING One cause of anode effects in every smelter is blocked feeder holes. If the hole is blocked, no alumina feeds the pot and an anode effect results. As mentioned in Section 1.2.12 and illustrated in Figure 1-19, some plants have developed sensors that detect if the plunger touches the bath by measuring a voltage across the feeder (52, 53). These require the plunger to get wet to operate. It has been established that it is not necessary to get the plunger wet to break a hole (Section 3.4 and 5.4.3). Many plants (and individual pots within most plants) have a zero plunger penetration in the bath. This stops the voltage sensing design from working. Dags on the plunger which cause extra resistance for the voltage signal, also stop the sensors from working. Trials of this type of cylinder at Dl 1 also had problems of arcing out some electrics if the crustbreak is extended into the bath during a high voltage anode effect. At the request of the author, Atlas Copco has designed a mechanical fitting in their cylinder that will sense the end of the stroke. The piston hits a button in the bottom block of the air cylinder. This sensor sends an air pressure signal to an electric switch on the top flange. This, in turn, tells the computer that the cylinder has reached the end Us of stroke. A Although end stroke sensing does not inform the operator/computer that the hole is "open" (as such), it does indicate the crustbreaker is fully extended. Provided the crust is not significantly below the plunger tip (which is usually the case), then the design CHAPTER 10 FEEDER DEVELOPMENT PAGE 368 will indicate blocked feeder holes. Liquid level control ensures that the total liquid level does not get too low. Usually, if the crust is low and the hole blocks, the material will build up over the hole and will quickly prevent the crustbreaker from reaching full stroke and the alarm is activated. Hence, this design is better than the designs that need to wet the plunger. At worst, there is only a minor delay before the blockage is alarmed. A timer is set at a limit above the normal stroke time of the cylinder. When this time is exceeded, an alarm is activated by the computer to tell the operator to check the pot. In this way a blocked feeder is detected immediately it happens. This minimises the time an operator is exposed to the hazards of unblocking holes and removes the cause of anode effects well before they occur. Another advantage of the design is a reduction in time that the plunger is wet. Instead of using a computer dwell time to control when the plunger is raised, the plunger is raised when it has positively proved to have travelled full stroke. For pots with multiple feeders per DCV, when all have reached full stroke, all are raised. For independent feeders, smarter logic can be used. For example, if a blockage is detected, the feeder can be turned off and the other feeders in the pot feed more frequently until the blockage is fixed. The crustbreaker can cycle several times to clear the blockage. Only when the blockage is not cleared after, say, 4 attempts, does an alarm inform the operator that the problem can not be fixed by the computer. This type of feature is a very positive way of addressing anode effects and the health issues associated with dags. Both the A3 feeder and split header pots have end stroke sensing facilities installed in their control logic and the A3 feeders have end stroke sensing facilities on the crustbreak cylinder itself. Trials are underway with end stroke sensing on AEDD, sequential feed and A3 feeder pots, however, to date the system has not been fully commissioned as the author was transferred from the Potrooms in mid 1994. It should CHAPTER 10 FEEDER DEVELOPMENT PAGE 369 only take a few days to commission it. Despite this lack of interest on use of this design feature, there have been no mechanical failure of cylinder components in almost 2 years of continuous operation of 20 cylinders. Although the electrical signals and alarms are not activated, the mechanical equipment has worked without fault. It is hoped that this installed design feature which has great potential will be fully commissioned and tested. 10.7 VALVE AND CYLINDER DESIGN These trials operate different types of cylinders and a mixture of different valves in the same pot to give an objective comparison of the different brands for identical pot conditions. This avoids pot-to-pot or time-related variables. Atlas Copco have a proven record for DCVs on many plants (Table I-IT). Mac have an equivalent DCV at half the cost. The top flange of the A3 and sequential feeders have provision to use both types interchangeably. Half the valves trialed were Atlas Copco and half were Mac. The Atlas Copco valves have had no faults in thefirst tw o years of operation, versus 8 failures for Mac valves. Four different brands of cylinders are being tested - Parker, Terry, Ortman and Kempe. In addition, Kempe have designed a different feed cylinder such that there is no alumina bypass on the dosing unit. The latter feeders were removed from operation by mid 1994 due to constant attention to pneumatic controls. Initially there were problems with most types from seals being too tight. One needs to accept a little air leakage at room temperature as seals tighten at elevated temperatures (Refer Section 8.4 on air leaks at room temperature versus that at operating temperature). Once this was understood and modifications made only two cylinders (Kempe and Terry) have failed in service after 20 months operation. This has been very good performance for such a radical design of cylinder. Parker and Ortman cylinders have never had any problems. Parker and Ortman feed cylinders and Atlas Copco valves should be used for future pots. CHAPTER 10 FEEDER DEVELOPMENT PAGE 370 10.8 INTEGRATED VERSUS INDEPENDENT FEEDERS Table 10-1 compares the main features of the different designs examined in this study. Table 10-HI summarises the previous references to integrated and independent feeders to verify conclusions on the relative benefits of each design. It has been established that independent feeders are superior to integrated feeders in almost all aspects except capital costs. It is clear that the independent feeders are more accurate, provide a more repeatable dose, have lower air usage and have potentially longer life. However, the capital cost is higher. The benefit on current efficiency and anode effects are subjective as no comparative studies have been published. Although nothing is published on pot benefits of independent feeders, dissolution studies show potentially better mixing by direct feed (Section 1.3.13) as well as the better shot size accuracy verified by Portland trials (Section 2.4.5). One would expect benefits to pot operations. What is unknown is if these are significant enough to justify three times the capital cost. It seems inconsistent of the reputation of companies such as Pechiney, Hydro Aluminium and Reynolds to install independent feeders with poor economic return for all their installations since 1979. Although no data are published, it appears that these companies believe the higher cost is justified. The current trials at Portland have the potential to identify the benefits of the different feeder designs. One of the features of the trials is that each pot can be changed to different feeder schemes by turning a switch. In this way, pot-to-pot differences do not affect comparisons. These comparisons may take several years to complete due to financial restrictions on the number of pots that can be trialled. Conclusions would come much quicker if more pots were brought on line. Results to date after 2 years operation show that the A3 has superior performance to normal AEDD feeder pots with respect to metal production, but have a higher than normal frequency of anode effects (Table 9-VI). CHAPTER 10 FEEDER DEVELOPMENT PAGE 371 TABLE 10 - m ADVANTAGES AND DISADVANTAGES OF INDEPENDENT FEEDERS FEATURE SECTION FOR 1 DONT NEED TO BREAK A HOLE EACH FEED 1.2.5, 1.2.6 2 LESS AIR USAGE 1.2.5, 1.2.6 3 REDUCED DUST LOSS BETWEEN CHUTE AND HOLE 1.2.5, 1.2.6 DIRECT FEED - BETTER DISSOLUTION 1.2.5, 1.2.6, 2.8 4 - LESS CRATER 2.8, 3.3(vii) 5 - HORIZONTAL VELOCITY VECTOR 2.8 6 LONGEST LIFE AIR CYLINDERS 1.2.15 LONGEST LIFE FEEDERS 1.2.15 7 - CONSISTENT FEED TO POT 1.2.15 8 - LOW MANTENANCE COSTS 1.2.15 8 SPOOL ANGLED TOP & BOTTOM - GOOD FOR SHOT SIZE 2.3.2 MORE ACCURATE SHOT SIZE S.D. 2.4.5, 2.6.4 10 -DCV HAS NO EFFECT ON SHOT SIZE 2.4.5 11 -DWELL TIME HAS NO EFFECT ON SHOT SIZE 2.6.4 12 DAG SCRAPER 2.8, 5.4.7 SLOW PISTON SPEEDS 3.4.1,8.3.4,8.6.1,8.8.12 13 -LOW FATIGUE 3.4.1,8.3.4 14 -LONGER SEAL LIFE 3.4.1,8.3.4 VALVE NEAR CYLINDER 3.4.1,4.6 15 -HIGH PRESSURE AT CRUST 3.4.1 16 -REPEATABLE DWELL TIME 4.6 17 -NOT AFFECTED BY CYLINDER VARIABILITY 4.6 18 LOW NUMBER OF ANODE EFFECTS PER DAY 3.5.2, 10.6 19 LESS EFFECT FOR MUFFLERS BLOCKING 3.5.3.3 LOW DWELL TIME AS NOT WAITING FOR ALUMINA 35.8,5.4.1:1 10 -LESS DAGS 3.5.8, 4.6 LARGE STROKE DISTANCE 6.5.6 12 - COOLER SEALS 6.5.6 13 - LESS PLUNGER PREHEAT 6.5.6 14 NO ORE LEAKS (44% OF INTEGRATED FEEDER FAILURES) 8.3, 8.4.1 NO SPRING 8.3, 8.8.4 15 - LESS JAMMING OF SPOOL 8.3, 8.8.4 16 - REPEATABLE SPOOL STROKE SPEED 8.3, 8.8.4 17 LESS ECCENTRICITY - LONGER SEAL LIFE 7.3.3:4, 8.4.4 18 NO ALUMINA NEAR SEALS - NO INGRESS 8.4 19 BETTER INSULATED 8.7, 8.7.2 20 NO NEED FOR TRACKING SYSTEM AS LOW FAILURE RATE 9.6 AGAINST 1 NO PREHEATING OF ALUMINA 1.2.5, 1.2.6 2 ABOUT 3 TIMES THE CAPITAL COST 1.2.5, 1.2.6 3 NO DIRECT FEED IF FEED EVERY 2-3 TIMES 2.8 4 DEEP PENETRATION - NEED COSTLY PLUNGERS 6.5.5 5 SOME SPOOL SHAFT FAILURES 8.8.12 6 USE VITON - NOT A HIGH TEMPERATURE RATING 7.33.2 CHAPTER 10 FEEDER DEVELOPMENT PAGE 372 This may be due to a lower than normal concentration of alumina in the bath. The time in anode effect suggests that the pot is on the edge of the optimum operating concentration thus assisting production rate. As there are only two A3 pots operating at present, one cannot make too many conclusions on comparative trials, however, to have both pots with metal production in the top 25% of pots is very encouraging. It is the author's opinion that the benefits of independent feeding would only have a cost benefit for large pots viz. 180kA or larger. This is due to the accurate dosing achievable on 1-2 feeder systems using a common DCV (as explained in Section 2.4). It is unlikely that the air savings and marginally improved dosing accuracy will make a difference to the pot economics on smaller pots. CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 373 CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS 11.1 SUMMARY The research conducted by this study has resulted in benefits to Portland and many other smelters world wide. The recommendations made by the author to Alcoa plants in 1992 and developments by teams at all Alcoa plants resulted in almost double the feeder life and half the cost between 1991 and 1994. Portland alone has reduced costs by $1.4 million per annum. Feeder development has been a major enabler to halving anode effects across Alcoa since 1991. This development also addressed safety and health improvements and involved operators and tradespersons in the implementation of design and procedure changes. CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 374 11.2 INTRODUCTION The impact of the improvements to existing feeder designs on feeder life, costs and anode effects are summarised in this Chapter. The importance of designs to assist the health of people who operate, change and service feeders is included. 11.3 PORTLAND RESULTS Figure 11-1 illustrates the changes made at Portland to improve the AEDD feeder. A more detailed history is in Appendix 9. In the next sections is an explanation of the results of this activity. 11.3.1 Plant Performance The objectives of the research into optimising feeders were three fold: (i) reduction in feeder changeouts (for health benefits), (ii) reduction in maintenance costs (for plant economic return), and (iii) reduction in anode effects (for global warming and pot stability reasons). Reduction in anode effects is the end result of the other two. Implementation of this research at Portland has had slower results at Portland than at several other plants despite using the same ideas of improvement. This has been due to the more difficult conditions at Portland as shown in Table 11-I. The combination of higher temperature, 5 feeders/DCV, deep penetration, poor cylinder design, grease breakdown and high kinetic energy resulted in a frustratingly slow change in performance. From mid 1993, a step change occurred in feeder failures (Figure 11-2). The spike in mid 1993 was a result of smaller shot size from the inserts and the poor choice of PA spec cast iron that actually reduced plunger life. a 1 2 o — to • • - < •J •' "• ' • -1 "•" 1 o s * 1 H s — o> LL 1 o> - w a Hi z Hi o HI 60 _ . . OH 4 H • — -> [ -1 2 < oo s LL O) -t a z. o CO < 1 -1 ^ S < CN 5 U- Oi -j o z o 1 a) T— < CJ -> -> 2 SPOOL S TERR Y T S WASHE R LURHAUL S K r s W POTS ) tr >}IU Pz o O CO < ccS < LU Z| m LU H CO CO 1 X fe to l-CO Q Oi Szl^Sco ^ rf *Q. TABLE 11 -1 PORTLAND VERSUS OTHER FEEDERS PORTLAND Vs OTHER INTEGRATED FEEDERS DIFFERENCE MECHANICAL FAULT CAUSE OF FEEDER FAILURE PORTLAND IS HOTTER (60°C AVERAGE) Cooked seals Ore leak Springs shrink Ore leak Plunger preheat Pointed Dags 5 FEEDERSA/ALVE NOT 2-3 FEEDERSA/ALVE Variable response Longer dwell time Pointed Dags Variable shot size Anode effects High air volume Air usage high Sensitive to faults Anode effects Localised high gas Dags Pointed ORIGINALLY VITON SEALS NOT TEFLON High wear rate Ore leaks Requires grease Grease breakdown Ore leaks Muffler blocks Blocked feeders Valve clogs Air leaks Blocked feeders PORTLAND Vs INDEPENDENT FEEDERS CAUSE OF FEEDER DIFFERENCE MECHANICAL FAULT FAILURE PORTLAND IS HOTTER (60°C) see above see above 5 FEEDERSA/ALVE NOT 1 FEEDERA/ALVE see above see above 2-3 TIMES THE CRUSTBREAK CYCLES 2-3 times lower life of seals Ore leaks 2-3times lower life of plungers Pointed 2-4 TIMES THE PLUNGER DWELL TIME Pointed CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 377 By late 1994, the failure rate had dropped from the 1993 average of 230/month to under 50/month. By mid 1994, feeder changeout rates were about half the record lowest month ever. With a normal life of about a year, it would have been expected that there would have been a jump in failures by mid 1994 due to the high changeout rates of mid 1993 when changeouts were accelerated to install spool inserts and remove the poor wear rates of PA spec cast iron. This did not happen. In fact, failures dropped to under 30/month which is the level required for a 5 year feeder. The changeout rate continued at about this level right through 1995. Thus, the improvements were working. However, graphs of frequency of failure may give a false impression of performance. For example, if all the feeders were rapidly replaced, then there would be a period immediately afterwards when the frequency of failure would be small; the "honeymoon period". In fact, the performance could be no different to normal. Similarly, failure rate analysis may be affected by the acceptance criteria of when to changeout feeders. For example,tf a supervisor decided to change the standard of pointed feeder from a minimum of 50mm O.D. to 40mm, then there would be a rush of changeouts that would have no bearing on changes made at repair prior to installation. CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 378 FIGURE 11 - 2 PORTLAND FEEDER CHANGEOUT HISTORY Feeder Repair Rate: C—Control Chart Date Repaired between 1 January 1991 and 31 December 1995 Component Typei (ALL) Iff Limits For n=1: 400 - £ "2 200 UCL=16S-7 il C= 144.7 LCU=12Qu6 1DD 1 | • • • i i i i i i i ' i ' ' ' ' ' I I I i ' ••••"• JMJ91 JUL91 Jfct|J9a. JUL92 JAW93 JUL95 JANS4 JUL94 JANS5 JUL95 Date Repaired FEEDER LIFE (YEARS) 1991 1992 1993 1994 1995 CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 379 A better analysis of the effect of changes in design or operation is to examine the life of feeders from the date installed, not removed. Figure 11-3 shows the life of 20%, 50% and 80% of feeders from date installed. This shows a steady median life at about 200 days through to 1991 before quality control actions were initiated. Life rose to 300 days in 1992 but dropped in 1993 back to the original life when the smaller shot size of inserts and poorer life of PA spec cast iron started to have effects. From mid 1993, life rose rapidly due to short stainless steel plungers, better grease and holes in assemblies (to stop jamming of the spool). Between 1992 and 1994, anode effects at Portland halved partly due to better hardware and troubleshooting that was developed by this investigation. A lot of work was done by others on computer control (e.g. Fethon Nahoum), operator attention to detail and a lift in awareness of the importance of solving excessive anode effects for improved pot control. Feeder improvement was significant, but was not the only cause. It was an enabler. Costs and changes dropped accordingly to record low levels. By the end of 1994, all capital and expense costs for this project have been covered by savings. Figure 11-4 shows the annual changes to frequency, costs and anode effects per day for Portland. 11.3.2 Improvements for Health Aluminium smelting pots have a range of hazards that can affect the people who attend them - heat, dust, fumes and falls (1,72,73). Many of the jobs are affected by strong magnetic fields that tend to move equipment away from the intended direction. The magnetic fields are so strong that they clear the magnetic signature on credit cards and can hold a 25kg crow bar in an almost vertical position. Spanners and wrenches can stick onto fixed equipment and two hands are required to release them. Thus, movement of tools is very awkward and can affect the safety of doing what appears to be simple jobs. 2 => £ -0 v »- i*-o .— a •_ w n > —I CO CO 3 co ^ ry UJ « O S x co E z co cog TO Eg 4 || 3 Q. LL to C O CD O LU > LU -J C > a. z c 2 < 3 O IO ? •- CL co 0 co o Q (V tn Q O0OOCOOO z K S 8 S K M •»- 0 _| .- H- * r- £ £) O *t IA X wy CO u. 0 H O t, 5 CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 381 FIGURE 11 - 4 PORTLAND FEEDER PERFORMANCE - SUMMARY OPERATING FEEDERS ANNUAL REPAIR COST (Aim) 1991 1992 1993 1994 10 NUMBER OF OVERHAULS COST PER OVERHAUL (AS) 0 -:—^—^t/Mm. ^_ FEEDER LIFE (YEARS) COST PER TONNE OF METAL (A$) ANODE EFFECTS PER DAY Note: (i) Anode effects per day "est 1995" is the estimate for 199S based on the average for Janua(y-June,1995. (ii) "Improved feeder life will support quantum leap improvements in anode effects, power consumption, and safety. In addition, longer feeder life offers the opportunity to substantially reduce maintenance costs." RE. Seymour, 16.4.92.(99) (iii) These "research activities have been a key enabler to achieving the reduction of anode effects per cell day in the international locations and, consequently, the significant reduction of perfluorocarbon emissions" R. R. Taybr,25.7.94.(100) CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 382 Throughout the development of improvements to the existing feeders or new feeders, the designs have been heavily influenced by the people who do the job - operators who changeout the feeders and tradespersons who overhaul them. Some of the items developed to assist people included better feeder racks and lifting slings. One example of a design that helped operators and tradespersons as well as improving feeder life and costs was the design of better mounting flange fittings. A common cause of feeder failure is jammed objects (Table 1-V). This is often from loose feeder bolts, nuts or washers. The real cause of this is the manner of installation and/or insulation breaking down. In addition, the nuts, bolts and washers are difficult to handle in the strong magnetic field, so dropping objects into the alumina in the hopper is not uncommon. The job of changing feeders is a hot, dusty, fumey operation, and operators/tradespersons are exposed to fall hazards from heights up to three metres. It is not a pleasant job. An alternative design was required that would (i) allow minimum time on the pot for changeout, (ii) be simple to modify existing feeders/superstructure, (iii) have less objects that may fall or have to be carried to the top of the superstructure, (iv) be non-magnetic if possible and (v) be cost effective, especially if 2,000 feeders were installed. Many mounting designs were trialled at Portland...Huck bolts, overcentre clamps, barrel clamps, G clamps. All these alternatives were expensive and had weight or fatigue problems. The author also wanted to find a way to get better venting of air from the feeder to prevent ore leaks that are caused from aeration of alumina and break down of the seals at the bottom of the dosing unit. Eventually a simple design was developed of a fixed insulation washer on the flange (at each of the four bolt holes) with boomerang brackets (Section 8.4.8 and Figure 8-4 and 8-5). CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 383 The result was a once-off modification that solved the installation problem at low cost. The time required on the pot has more than halved and the number of parts reduced to 2 brackets and two bolts (which had the washers fixed to them) instead of the 28 parts in the old design. The $50/flange cost ended up saving about $70,000 on parts per year as almost all items are now reused. Not only has the cost reduced, but the big attraction was that the design allowed any air leaks from the cylinder to vent from the feeder without causing ore leaks. Ore leaks are the most common cause of failures of integrated feeders (Table 1-V) and a contributing cause of failures for independent feeders. As explained in Section 8.3.1, the main cause of ore leaks is a leak from the rod seal of the cylinder aerating the alumina nearby and causing alumina (ore) to fall into the pot unmetered. Since the boomerang brackets and fixed insulation was installed at Portland, ore leaks have almost disappeared (Figure 9- 4). Now only excessive cooking of seals causes ore leaks. As well as the change to the flange fittings, non-magnetic lifting slings and racks to hold feeders safely in a vertical position assisted the operators from having to force tools against the magnetic forces and gave them a stable place to store and remove new feeders without any hazards. These racks also were used to store all nuts, bolts and tools for the changeover job,. The racks contained covered boxes to store installation and removal tags for the tracking system. Thus, all items required to change feeders were located in easy reach of the person doing the job. Hence, astute asessment of the process and people's needs have solved hazards as well as saving costs. 11.3.3 Worker Involvement In this project, operators and tradespersons have been directly involved both in development and in normal maintenance. Operators and tradespersons have unlimited access to all screens in the computer tracking system, so they take an interest in the history of feeders. CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 384 Tradespersons were frequently consulted on how things were going and what they thought of any trials. They often identified opportunities and were a key indication of trends before data were sufficient to identify the trends. Although they were not always correct, it was useful data to keep management and engineers on track. They also appreciate that they have ownership of, and are a valuable input into the process and are rightly proud of improvements. It is interesting that, even when their jobs were at some risk as failures dropped, there were only positive statements from tradespersons of how well "we" have done. It was a team effort. This is the only way to get sustained improvement. Modifications for health developments have been warmly received by operators and tradespersons. They can see their job is safer and easier. This also reflects in a more positive attitude to a job that is often boring, hot, dusty and fumey. As operators can input why they removed the feeder, they know they have an influence on the process. The success of feeder development at Portland (and other smelters that have followed Portland's conclusions and recommendations) has been largely affected by the involvement by these operators and tradespersons. Their contribution cannot be under- emphasised. 11.2 ACTIONS AT ALCOA SMELTERS As discussed in Section 1.3.4, in 1992 the Alcoa Primary Metals Quality Manager Bob Seymour (who was in charge of all international and USA smelter quality improvement teams) directed that each plant follow the advice of the author on feeder improvements. His letter to all Smelting Managers is in Appendix 4 (99). Following the Kissane Feeder Recommendations (Table 1-VH), action was taken at all Alcoa plants on some or all of these items. Figure 11-5 shows 1991 to 1994 data from all Alcoa plants, and Table 11-11 shows highlights from these plants. Across Alcoa there was a reduction of 44% in changeouts and reduction of 48% in costs....over $2m savings. Anode effects dropped by 70% (Figure 11-6). CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 385 FIGURE 11-5 ALCOA POT FEEDER PERFORMANCE - BY PLANT OPERATING FEEDERS ANNUAL REPAIR COST (A$m) NUMBER OF OVERHAULS COST PER OVERHAUL (A$) B1991 • 1993 • 1993 0199* • 1994 • 1995 ill 9 Jt- HCL FEEDER UFE (YEARS) COST PER TONNE OF METAL (A?) - Q1994 Tl—' a 1995 — 1 (__ •£••'••}• - 1-- .zn( CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 386 TABLE 11-H COMMENTS FROM ALCOA PLANTS ON FEEDER DEVELOPMENT PORTLAND - increased wear by 30% when reduced shot size - dropped plunger life by 40% on change to P.A. cast plunger - plunger life 5 times better for short 310 cast stainless and 15 times better for Nicrofer - increased frequency of changeouts to address A.E's - not changing feeders on off line pots - air leak rate average is FIGURE 11 - 6 ALCOA FEEDER PERFORMANCE - SUMMARY OPERATING FEEDERS ANNUAL REPAIR COST (ASm) NUMBER OF OVERHAULS COST PER OVERHAUL (AJ) Note: (I) Anode effects per day "est 1995" is the estimate for 1995 based on the average for January-Jun«U995. (il) "Improved feeder life will support quantum leap improvements in anode effects, power consumption. _ and safety. In addition, longer feeder life offers the opportunity to substantially reduce maintenance cost* R E. Seymour, io.4.a£.v*'/ (iii) These "research activities have been a key enabler to achieving the reduction of anode effects per celday in the international locations and, consequently, the significant reduction of perfluorocarbonOTissons^ ^ CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 388 The results of some plants were comparable to Portland. T7 for example, reduced turnover by 82% and costs by 90%. T16 had the highest life at about 5 years. Most of the plants that had poorer performance in 1991 (Portland, T6 and T9) had to invest in a rebuilding campaign to obtain control of performance, but showed evidence of the projected benefits by early 1995. Less exposure of operators and tradespersons to dust, heat, fumes and potential fall hazards has been a pleasing benefit. A very important feature of the investigations has been the extensive involvement of tradespersons and operators. This had a significant impact on results. A better understanding of what causes anode effects has resulted. This has reduced the variability of feeding so that computer changes can fine tune the pots. It is considered that feeder improvements were contributors to the reduction in anode effects across Alcoa according to Dick Taylor (Director of Technology, Primary Metals, Alcoa) (Appendix 5). Taylor also verified the impact of this investigation on the "improvement in the design, control and functionality of the alumina feeder" and that these "research activities have been a key enabler to achieving the reduction of anode effects per cell day in the international locations and, consequently, the significant reduction of perfluorocarbon emissions" (100). Experience has shown that two of the 10 recommendations did not achieve what was expected. Spool inserts proved to be beneficial on the large feeders at Portland, but were not as significant at smaller plants. The PA spec cast iron plungers (used at Portland and T6) proved to give worse life than the previous HR cast iron. Short 310SS plungers were recommended to all plants in mid 1993 but Portland is the only Alcoa plant which has changed to non-cast iron plungers to date, despite limited trials in T6 andT9. CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 389 The other eight original recommendations proved very successful. All plants found that blocked feeder holes were mainly due to covering after set, not plunger speed. Dwell time was halved at T15 and T9. D12 saved 250cfm of air by addressing leaking DCVs. Portland, T16 and T7 used their tracking systems and procedures in a very proactive manner and identified many problems that were never appreciated before. Except for Portland and T6, most plants did not proceed with dual dwell time due to computer restrictions. Clearly, not all the results are due to the Recommendations per se, but concentrating on feeder hardware (which had not been done as intensively before) showed many ways that the plants could be improved. The key to any further improvements and cost cutting is an effective tracking system. Not many plants had this area under control. By 1995, most plants had difficulty in even doing a Pareto chart on why feeders were removed (though most had lists for what was done in the workshop). Portland has found that the pot conditions are usually the major reasons for failure...not the way the feeder was repaired. It is important that the cause is traced and a good tracking system is required to do this. There is still considerable scope for plants to make significant savings in feeder maintenance. Installation of stainless steel plungers, use of lubrication, air leak venting, using breathing holes in the assembly and addressing optimum plunger depth show great promise in many plants (both Alcoa and non-Alcoa). Note that use of cast plungers is unwise from an economic point of view and cast plungers have potential to create dags initially due to imperfections in surface finish. Any plunger trials should be comparative trials against a control plunger in the same pot, for the same time and starting from the same original diameter. Although a lot of progress has been made, there is still a wide gap between the 2 year Alcoa feeder life and the 5-6 years average claimed by most independent feeder plants. CHAPTER 11 IMPACT OF RESEARCH ON ALCOA AND OTHER SMELTERS PAGE 390 11.3 ACTIONS AT NON-ALCOA SMELTERS Close contact with other smelters has resulted in many actions which have been significantly affected by this research. Some items are listed in Table 11-IH. TABLE 11-ID ACTIONS AT NON-ALCOA PLANTS T2 universal joint for reduced misalignment between piston rod and plunger shaft T3 alternative plungers independent feeder trials T7 change in piston rod shape to stop fatigue trials on slow plunger speed to stop fatigue co-operative trial with Portland of alternative seals T8 retrofit of Terry FEC in Parker cylinders T10 implementation of Terry cylinders implementation of Terry FEC changed from routine feeder changeout to repair on failure D3 alternative plungers do not break every feed D11 implementation of Terry cylinders In addition to the above, Terry Fluids are installing their cylinders at Capral (formerly Alcan) at Kurri Kurri, NZAS (Comalco in New Zealand), Comalco Bell Bay, Boyne Island and Dubai (United Arab Emirates) as a result of the seal research at Portland. The cost/benefits and extent of life improvements are not available, but comments by representatives of these plants testify to significant improvement in feeder life. CHAPTER 12 SUMMARY OF RESULTS PAGE 391 CHAPTER 12 SUMMARY OF RESULTS 12.1 RESEARCH OBJECTIVES From early 1990, this research into feeders has considered two basic objectives (Section 1.3.1): (i) optimise existing feeder designs (especially the AEDD feeder), (a) reduce failures (b) reduce anode effects (c) reduce costs (ii) compare integrated and independent feeder/crustbreaker units and develop a state-of-the-art feeder which has the following features. (a) life of over 5 years, (b) cost effective in both capital and maintenance costs, and (c) can be retrofitted in integrated pots easily. Consider the results of this research against these objectives. 12.1.1 Feeder and Pot Performance Chapter 11 illustrated that, compared with a 3 year average from 1991-1993, in 1995 the Portland results were as follows (Figure 11-4): (a) feeder life increased by 83%, (b) anode effects reduced by 68%, and (c) feeder repair costs reduced by 82% or $1.4m per annum. CHAPTER 12 SUMMARY OF RESULTS PAGE 392 From 1991 to 1994, Alcoa results were as follows (Figure 11-6): (a) feeder life increased by 52% (b) anode effects reduced by 72%, and (c) feeder repair costs reduced by 48% or $3.4m per annum. Note that the reduction of anode effects was not solely due to feeder life/performance improvement but to many actions taken by others. Feeder research was a significant enabler to this improvement (as per Taylor note (Appendix 5)). About 10% of the cost reduction was due to the reduction of feeders on line because of the global reduction in aluminium manufacture in 1994-1995. However, the costs in 1995 are considered to be much lower than in 1994 as a result of implementation of major changes that inflated costs at several smelters in 1994. Plant benefits are the result of sustained efforts by many at each plant. This research helped direct actions to the most appropriate areas. Although data is not available, it is believed that there have been significant improvements to non-Alcoa smelters because of this research. 12.1.2 Feeder Development Compilation of information in literature, data and comments from plant contacts and data developed in this study (Table 10-111) showed that the independent feeder is superior to the integrated feeder in almost all aspects except capital cost. The development of the A3 and Sequential Feeder designs achieved most of the benefits of the commercially available independent feeders yet at reduced costs and they could be retrofitted into integrated feeder plants on operating pots. Thus, the objectives of capital cost and ability to retrofit were attained, but it is still too early at the time of writing to identify if life and maintenance costs are adequate for these designs. Until a larger sample size of test pots is placed on line, it will not be possible to gauge feeder life and the effect on pot performance for these designs. CHAPTER 12 SUMMARY OF RESULTS PAGE 393 12.1.3 Conclusions on Achievements Against Objectives The objectives of this project were met with respect to optimising feeders...reduced failure rates, reduced anode effects, reduced repair costs. Feeder designs were developed that have the potential to be cost effective and can be retrofitted in operating pots. Thus, the objectives originally set were largely achieved. 12.2 FUTURE WORK There are several ideas that can achieve cost savings as well as improve life of existing feeder/crustbreaker units in all smelters (whether integrated or independent): (i) stainless or alloy plungers (ii) small penetration in the bath (iii) reuse feeders from off-line pots (iv) special attention to repeaters and feeders with high cost repairs (v) air leak tests on overhaul (vi) air audits in potrooms Optimisation of the integrated feeder at Portland is basically on "hold" from 1994 as the recent changes to grease, guides, seals, bush wiper, inserts and plungers take effect. Most of the above points are long term design changes, so time is required to collect enough data to measure effectiveness before further changes are made. The main direction of future development of feeder schemes at Portland will be the testing of the new feeder schemes; sequential feed, A3 feeders and end stroke sensing. The second stage of an additional ten A3 pots rests largely on the results of these trials and the availability of money in the current poor metal price situation. There is considerable interest by operators and technical personnel in these ideas, but these will only progress if there is a clear picture of enhanced pot efficiency. Unfortunately, this is difficult to see unless a large number of pots is operational, due to the large variability in pot performance. CHAPTER 12 SUMMARY OF RESULTS PAGE 394 At other Alcoa smelters, there are now teams of tradespersons, engineers and supervisory people continually optimising their feeders. Trials or implementation of many items raised in this report are continuing at other non-Alcoa plants, especially for cost control in the current climate of poor profitability of aluminium production. 12.3 CONCLUDING REMARKS One item that Bob Seymour (99) mentions in his note to the Alcoa Smelting Managers in 1992 should be stressed in assessing the results of feeder optimisation work. "Improved feeder life will support quantum leap improvements in anode effects, power consumption, and safety. In addition, longer feeder life offers the opportunity to substantially reduce maintenance costs." This needs to be stressed. The main target for this work in smelters is not a maintenance activity, though a lot of the work to achieve the target improvements is intimately involved in feeder repairs and feeder design. It is very easy to measure the priorities, success and failure of feeder research on reduction in maintenance costs. 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Knapp, L., (Boyne Smelters Limited) discussions at Fifth Australian Aluminium Smelter Technology Workshop, Sydney, 27.10.95. 103. Purdie, J., (Comalco), "Effect of Feeder Design on Alumina Concentration Control and Associated Process Variation", Fifth Australian Aluminium Smelter Technology Workshop, Sydney, 27.10.95, pp628-650. APPENDICES Appendix 1 Examples of Shot Size Control Charts Page 1 APPENDIX 1 EXAMPLE OF SHOT SIZE CONTROL CHARTS 1520 ^1500 ^1480 N 5)1460 O1440 x W1420 1400 1400 12 3 4 5 6 7 8 9 101112131415161718192021222324 25 SHOT NUMBER 1 3 5 7 9 11 13 15 17 19 21 23 0. 800 o 600 400 200 UJ > 0 UJ 1 3 5 7 9 11 13 15 17 19 21 23 25 SHOT NUMBER APPENDIX 2 OFF-SITE SHOT SIZE DATA Appendix 2 Off-site Shot Size Test Data Page 1 EFFECT OF ALUMINA TYPE, FEEDER DESIGN, INSERT AND SPOOL SHAPE RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (s) ] TYPE TIME (s) 0.6 1.9 1.6 2.0 2.S 3.0 3.5 4.0 46 5.0 10-Jan 1 STANDARD PARKER FINE KEMPE 175mm + 1.0/1.1 AVERAGE 878 890 1063 1195 1372 1483 1557 30mm SPACER STDOEV 4 8.5 9 10 a.5 a.s 8.5 10-Jan 2 PA#1 53mm INSERT PARKER FINE KEMPE 175mm* 1.0/1.1 AVERAGE 1398 1374 1384 1442 1445 1435 1423 (CAST IRON) 30mm SPACER STD DEV 1.5 1.5 2.5 1 1 2 1 3 10-J an PA#1 53mm INSERT PARKER FINE KEMPE 175mm* 1.0/1.1 AVERAGE 1414 1448 1464 1470 1447 1424 1455 (ENAMEL) 30mm SPACER STD DEV 2 1.5 2.5 1 1.5 1 5 10-Jan 4 A2 STANDARD PARKER FINE KEMPE 175mm + 1.0/1.1 AVERAGE 690 856 852 853 852 854 85B 30mm SPACER STD DEV 8 1.5 1.5 1 1.5 1.5 1 5 10-Jan TS#1 STANDARD TERRY FINE STAINLESS ST. 0.6/0.9 AVERAGE 1498 1594 1638 240mm STDOEV 1.5 1 1.5 10-Jan PA#1 BOLT HOLES PARKER e FINE KEMPE 175mm + 1.0/1.1 AVERAGE 742 955 1150 1394 1451 1510 1613 BLOCKED 30mm SPACER STD DEV 3.5 3.5 6.5 6.5 4.5 2.5 3.5 7 10-Jan PA#1 STANDARD PARKER NORMAL KEMPE 175mm* 1.0/1.1 AVERAGE 823 962 1112 1298 1408 1442 1527 30mm SPACER STD DEV 3 3 4 7.5 4.5 3.5 3 8 10-Jan PAH 53mm INSERT PARKER NORMAL KEMPE 175mm* 1.0/1.1 AVERAGE 1431 1471 1486 1489 1484 1471 1490 (CAST IRON) 30mm SPACER STD DEV 3 1 1 1 1 0.5 0.5 9 10-Jan PA#1 53mm INSERT PARKER NORMAL KEMPE 175 mm + 1.0/1.1 AVERAGE 1380 1491 1444 1462 1474 1510 1472 (ENAMEL) 30mm SPACER STD DEV 3.5 1.5 1 1 1 1 15 to 10-Jan A2 STANDARD PARKER NORMAL KEMPE 175mm + 1.0/1.1 AVERAGE 509 855 864 856 857 862 854 30mm SPACER STD DEV 16 15 0.5 1 1 1 75 11 10-Jan T6#1 STANDARD TERRY NORMAL STAINLESS ST. 0.6/0.9 AVERAGE 1403 1554 1808 1639 1664 1674 1685 240mm STD DEV 1.5 1 1 1 0.5 0.5 0.5 12 10-Jan PA#1 BOLT HOLES PARKER NORMAL KEMPE 175mm + 1.0/1.1 AVERAGE 735 952 1142 1319 1502 1560 1708 BLOCKED 30mm SPACER STD OEV 2 4.5 4 4 3.5 3 3.5 13 10-Jan PA#1 STANDARD PARKER UNREACTEO KEMPE 175mm + 1.0/1.1 AVERAGE 716 980 1182 1288 1419 1554 1687 30mm SPACER STD DEV 2.5 4 6 4 5.5 2.5 3 14 10-Jan PA#1 53mm INSERT PARKER UNREACTED! KEMPE 175mm * 1.0/1.1 AVERAGE 967 1428 1431 1431 1440 1450 1436 (CAST IRON) 30mm SPACER STD DEV 18 1.5 1.5 0.5 2 1 1 15 10-Jan PA#1 53mm INSERT PARKER UNREACTED KEMPE 175mm + 1.0/1.1 AVERAGE 1388 1391 1405 1416 1433 1428 1392 (ENAMEL) 30mm SPACER STD DEV 2 1 0.5 1 0.5 0.5 1 IS 10-Jan A2 STANDARD PARKER UNREACTED KEMPE 175mm + 1.0/1.1 AVERAGE 528 801 806 837 834 824 826 30mm SPACER STD DEV 14 2.5 2 2 1 3 1 17 10-Jan T6#1 STANDARD TERRY UNREACTED STAINLESS ST. 0.6/0.9 AVERAGE 1324 1471 1539 1585 1606 1626 1633 240mm STD OEV 3 1 1 0.5 1 0.5 0.5 18 10-Jan PA#1 BOLT HOLES PARKER UNREACTED; KEMPE 175mm* 1.0/1.1 AVERAGE 716 1149 1128 1435 1405 1584 1671 BLOCKED 30mm SPACER STD DEV 2.5 4.5 6 5.5 5.5 2 1.5 EFFECT OF SECOND FEEDER (SIZE.STROKE TIME.CYLINDER) RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (s) TYPE TIME(s) o.s 1.0 1.6 2.0 2.6 3.0 3.S 4.0 4.6 6.0 19 29-Jan PA #2 STANDARD PARKER FINE KEMPE 175mm* 1.0/1.1 AVERAGE 782 I 1001 1191 1381 1404 1555 1584 1670 1712 1715 30mm SPACER STD DEV 2.5 5 4 4.5 5 3 3 2 1 1.5 20 02-Feb T6#2 STANDARD TERRY FINE STAINLESS ST. 1.6/1.9 AVERAGE 1031 1426 1504 1584 1607 1633 1857 1686 1692 1703 240mm STDOEV 3 2 2 1.5 1 1 0.5 1 0.5 1.5 21 29-Jan PA #2 STAN0ARD PARKER FINE KEMPE 175mm* 1.6/1.9 AVERAGE 677 979 1205 1405 1530 1569 1604 1660 1676 1720 30mm SPACER STDOEV 5 3 5.5 4 5 2.5 2 1.5 1 2 22 29-Jan PA #2 STANDARD PARKER NORMAL KEMPE 175mm* 1.0/1.1 AVERAGE 795 1052 1323 1491 1566 1633 1679 1753 1791 1793 30mm SPACER STD DEV 2 3 3.5 3.5 2 2 2 1 1 1 23 31-Jan T8#2 STANDARD TERRY NORMAL STAINLESS ST. 1.6/1.9 AVERAGE 1144 1564 1631 1673 1692 1699 1694 1726 1729 1727 240mm STD DEV 3 1 1 0.5 1 0.5 0.5 1 0.5 0,5 24 10-Feb PA #2 53mm INSERT PARKER FINE KEMPE 175mm* 1.6/1.9 AVERAGE 560 940 1235 1280 1300 1340 1320 (CAST IRON) 30mm SPACER STD DEV 25 10-Feb PA #2 53mm INSERT PARKER NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 608 1033 1326 1348 1358 1361 1361 (CAST (RON) 30mm SPACER STD DEV 4.5 2 1.5 0.5 0.5 0 0 26 11-Feb PA #2 53mm INSERT PARKER NORMAL KEMPE 175mm* 1.3/1.5 AVERAGE 584 1074 1298 1344 1362 1364 1374 (CAST IRON) 30mm SPACER STDOEV 2.5 2 1 3 1.5 2 4 27 12-Feb PA #2 53mm INSERT PARKER FINE KEMPE 175mm + 1.3/1.5 AVERAGE 477 957 1245 1287 1305 1353 1384 (CAST IRON) 30mm SPACER STD DEV 2 2 1 1 0.5 0.5 0.5 EFFECT OF SIZE ON 81mm STROKE RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME |s) TYPE TIME (s) 0.5 1.0 1.6 2.0 2.5 3.0 3.6 4.0 4.6 6.0 28 14-Feb PA #2 81mm PARKER NORMAL KEMPE 175mm* 1.3/1.5 AVERAGE 1324 1944 2100 2119 2120 2119 2127 (NO INSERT) 30mm SPACER STD DEV 11 36 0.5 1 0.5 0.5 1 29 13-Feb PA«2 81mm PARKER FINE KEMPE 175mm + 1.3/1.5 AVERAGE 1302 1245 2048 2080 2097 2083 2080 (NO INSERT) 30mm SPACER STD DEV 4 2 2 0.5 0.5 0.5 0.5 EFFECT OF STROKE TIME AND DIFFERENT CYLINDERS RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (a) 3.5 4.0 4.6 6.0 TYPE TIME (3) 0.6 1.0 1.8 2.0 2.5 3.0 '3d- 28-Feb PA #2 STANDARD PARKER NORMAL KEMPE 175mm * 1.7/1.9 AVERAGE 823 1069 1438 1618 1745 1843 1724 30mm SPACER STDOEV 3.5 2 3 1.5 1 0.5 0.5 31 25-Feb PA #2 STANDARD PARKER NORMAL KEMPE 175mm* 1.2/1.7 AVERAGE 821 1174 1526 1736 1687 1833 1848 30mm SPACER STD DEV 2.5 2 1.5 1.5 0.5 1 0.5 32 28-Feb PA #2 STANDARD TERRY NORMAL KEMPE 175mm + 1.7/1.9 AVERAGE 922 1602 1817 1851 1867 1B73 1869 30mm SPACER STD DEV 2.5 1 0.5 0.5 0.5 0.5 0.5 33 28-Feb PA #2 STANDARD TERRY NORMAL KEMPE 175mm* 1.2/1.7 AVERAGE 992 1659 1818 1834 1852 1850 1850 30mm SPACER STD DEV 4.5 0.5 0.5 0.5 0.5 0.6 1 1852 34 24-Feb PA #2 STANDARD TERRY NORMAL KEMPE 175mm* 1.5/1.7 AVERAGE 705 1084 1481 1703 1756 1669 1 30mm SPACER STD DEV 2 2.5 2 1 0.5 0.5 Appendix 2 Off-site Shot Size Test Data EFFECT OF SPOOL STROKE (NO INSERT) WITH DIFFERENT STROKE TIME RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (s) TYPE TIME (S) 35 20-Feb PA #2 HIGH SPOOL TERRY NORMAL KEMPE 175mm* 1.3/1.5 AVERAGE 1802 2085 2075 2079 2069 2079 2071 80mm 30mm SPACER STDOEV 1 0 5 0.5 0.5 0.5 05 36 15-Feb PA #2 HISBWKM. TERRY NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 1186 1809 2180 2201 2202 2196 2189 80mm 30mm SPACER STD DEV 8.5 1 1 0.5 1 37 19-Feb PA #2 HIGH SPOOL TERRY NORMAL KEMPE 175mm* 1.3/1.5 AVERAGE 1667 1974 2044 2071 2066 2069 2073 80mm 30mm SPACER STDOEV 6 0,5 0.5 0.5 0.5 0.5 0.5 38 17-Feb PA #2 HIGH SPOOL TERRY NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 1463 2011 2068 2052 2110 2058 2062 80mm 30mm SPACER STDOEV 5.5 6.5 1 1 1.5 0.5 0.5 39 15-Feb PA #2 HIGH SPOOL PARKER NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 1224 1677 2061 1174 2166 2172 2170 80mm 30mm SPACER STDOEV EFFECT OF DIFFERENT FEEDERS FROM OFF LINE POTS DATE FEEDER RUN DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (a) TYPE TIME (a) 0.6 1.0 1.6 2.0 2.6 3.0 3.6 4.0 4.6 6.0 04-Mar PA #2 40 *1029 PARKER NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 856 1253 1668 1793 1819 1822 1866 30mm SPACER STDOEV 3.2 1.5 1.2 1.4 0.6 0.4 2.2 41 04-Mar PA #2 #1975 PARKER NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 748 1093 1519 1678 1728 1765 1803 30mm SPACER STD DEV 2.2 2.6 1.6 0.95 0.6 0.6 0.5 42 06-Mar PA #2 #1282 PARKER NORMAL KEMPE 175mm + 1.6/1.9 AVERAGE 789 1136 1531 1694 1780 1807 1801 30mm SPACER STD DEV 2.4 2.3 2.2 0.8 1.2 0.5 1.5 43 OS-Mar PA #2 *590 TERRY NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 739 1083 1434 1556 1659 1652 1683 30mm SPACER STD DEV 3.3 2 1.4 0.7 2.5 0.4 0.5 44 16-Mar PA #2 #474 TERRY NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 786 1235 1614 1690 1722 1743 1772 30mm SPACER STDOEV 2.8 1.7 0.8 0.4 0.4 0.6 2.7 45 10-Mar PA #2 #650 TERRY NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 940 1667 1622 1847 1848 1829 1844 30mm SPACER STD DEV 3.2 1.1 0.4 1.4 0.4 0.5 0.4 EFFECT OF DIFFERENT CYLINDERS WITH INSERT RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (a) TYPE TIME (s) 0.5 1.0 1.5 2.0 2.6 3.0 3.6 4.0 4.6 6.0 46 19-Mar PA #2 53mm INSERT PARKER NORMAL KEMPE 175mm* 1.7/1.9 AVERAGE 724 1217 1402 1437 1451 1430 1427 (STEEL) 30mm SPACER STD DEV 3 2.8 0.6 0.5 0.3 1 1.1 47 20-Mar PA #2 53mm INSERT PARKER FINE KEMPE 175mm* 1.7/1.9 AVERAGE 713 1131 1350 1399 1420 1445 1422 (STEEL) 30mm SPACER STD DEV 6.4 2.2 1.7 0.9 0.4 0.9 1.8 48 21-Mar PA #2 53mm INSERT TERRY NORMAL KEMPE 175mm* 1.7/1.9 AVERAGE 400 987 1288 1327 1346 1349 1342 (STEEL) 30mm SPACER STD DEV 20.5 2.7 1.6 0.6 0.4 0.8 0.8 49 21-Mar PA #2 53mm INSERT TERRY FINE KEMPE 175mm* 1.7/1.9 AVERAGE 609 1054 1280 1334 1352 1354 1364 (STEEL) 30mm SPACER STDOEV 2.2 3.9 1.6 0.5 0.3 0.3 0.2 EFFECT OF STROKE TIME.SIZE AND INSERT ON T6 FEEDERS RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (s) TYPE TIME (a) 0.6 I 1.0 1.6 2.0j 2.6 3.0 3.5 4.0 4.6 6.0 50 17-Mar T6#2 STANDARD TERRY NORMAL STAINLESS ST. 0.6/0.9 AVERAGE 1365 1646 1667 1680 1684 1682 1682 240mm STDOEV 2.4 0.6 0.4 0.4 0.3 0.2 0.6 51 18-Mar T8#2 STANDARD TERRY NORMAL STAINLESS ST. 1.1/1.3 AVERAGE 1251 1651 1683 1693 1692 1691 1689 240mm STD OEV 2.3 0.4 0.2 0.4 0.2 0.2 0.2 52 17-Mar T6#2 STANDARD TERRY FINE STAINLESS ST. 0.6/0.9 AVERAGE 1396 1609 1717 1677 1774 1783 1786 240mm STD DEV 6.4 1.1 3.9 1 1.4 1.2 1.2 53 18-Mar T6#2 STANDARD TERRY FINE STAINLESS ST. 1.1/1.3 AVERAGE 1056 1628 1663 1679 1697 1680 1660 240mm STD DEV 2.9 0.6 1.2 0.7 1.2 0.7 0.7 54 22-Mar T6#2 53mm INSERT TERRY NORMAL STAINLESS ST. 0.6/0.9 AVERAGE 966 1284 1287 1291 1288 1284 1284 240mm STDOEV 9.8 0.4 0.4 0.4 0.4 0.3 0.4 55 23-Mar T6#2 53mm INSERT TERRY FINE STAINLESS ST. 0.6/0.9 AVERAGE 1024 1292 1333 1281 1287 1291 1296 240mm STDOEV 22 0.4 0.4 0.4 0.5 0.6 0.7 EFFECT OF DIFFERENT FEEDERS. SIZE. AND STROKE TIME ON T11 FEEDERS RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME HI TYPE TIME |s) 0.5 1.0 1.S 2.0 2.5 J.O 3.6 4.0 4.6 6.0 56 24-Mar T11 #1 STANDARD LINDBERG NORMAL STAINLESS ST. 0.8/0.8 AVERAGE 1103 1565 1630 1653 1659 1673 1673 240mm STD DEV 5.6 0.6 0.4 0.4 0.4 0.4 0.4 57 23-Mar T11#1 STANDARD LINDBERG FINE STAINLESS ST. 0.8/0.8 AVERAGE 1078 1494 1554 1577 1594 1604 1612 240mm STDOEV 9 0.5 0.3 0.4 0.4 0.3 0.3 58 24-Mar Til #2 STANDARD LINDBERG NORMAL STAINLESS ST. 0.8/0.8 AVERAGE 966 1100 1247 1366 1459 1519 240mm STD DEV 2.4 1.8 2.4 1.6 1 0.6 1619 59 01-Apr T11#2 STANDARD LINDBERG FINE STAINLESS ST. 0 8/08 AVERAGE 905 1406 1493 1563 1575 1600 240mm STD DEV 6.9 1.2 0.4 0.4 0.6 0.8 1.2 1550 60 31-Mar T11 #2 STANDARD LINDBERG NORMAL STAINLESS ST. 1.3/1.3 AVERAGE 684 990 1215 1340 1429 1498 240mm STDOEV 3.4 2 2 1.5 1.2 1.6 0.8 1494 1530 81 31-Mar T11#2 STANDARD LtNOBERG FINE STAINLESS ST. 1.3/1.3 AVERAGE 696 963 1151 1318 1426 240mm STD DEV 3.2 2.5 2.5 1.2 1.2 1.6 1 mmm EFFECT OF SIZE ON T11 WITH INSERT DWELL TIME ll) RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA TYPE TIME (s) 0.5 1.0 1.6 2.0 2.5 3.0 3.5 4.0 4.5 5.0 1324 1324 1326 62B 02-Apr T11 #2 53mm INSERT LINDBERG NORMAL STAINLESS ST. 0.8/0.8 AVERAGE 695 1178 12S1 1304 0.3 240mm STD DEV 2.7 2.8 0.4 0.2 0.4 0.4 1319 1327 82A 02-Apr Til #2 53mm INSERT LINDBERG FINE STAINLESS ST. 0.8/0.3 AVERAGE 607 1141 1266 1303 1311 0.4 240mm STD DEV 2.3 2.2 0.4 0.4 0.4 0.2 Appendix 2 Off-site Shot Size Test Data Page 3 EFFECT OF STROKE TIME AND CYLINDER ON A2 FEEDFR RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME Is) TYPE TIMEIS) S3 25-Mar A2 STANDARD PARKER NORMAL KEMPE 175mm + 1.2/1.4 AVERAGE 497 8S6 839 849 860 843 857 30mm SPACER STDOEV 14 0.5 0.3 0.9 64 23-Mar A2 STANDARD PARKER NORMAL KEMPE 175mm* 1.7/1.9 AVERAGE 437 893 857 858 897 857 846 30mm SPACER STDOEV 1 0.5 0.3 0.5 65 01-Apr A2 STANDARD TERRY NORMAL KEMPE 175mm + 1.7/1.9 AVERAGE 407 915 912 909 915 917 918 30mm SPACER STDOEV 0.9 0.5 0.4 0.6 68 04-Apr A2 STANDARD TERRY NORMAL KEMPE 175mm* 1.2/1.4 AVERAGE 652 910 910 914 911 912 908 30mm SPACER STD DEV 1.8 0.3 0.4 67 04-Apr A2 STANDARD TERRY FINE KEMPE 175mm* 1.7/1.9 AVERAGE 506 936 933 938 942 946 940 30mm SPACER STD DEV 1.1 0.4 0.4 0.6 0.5 0.8 EFFECT OF SPOOL SHAPE RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (af TYPE TIME (SI 0.6 j 1.0 1.6 2.0 2.6 3.0 16 4.0 4.6 6.0 68 24-Apr PA #2 SHORT SPOO PARKER NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 356 309 1220 1609 Hi ,9 30mm SPACER STD DEV 1,3 1 0.6 0.5 69 24-Apr PA #2 81mm INSERT PARKER NORMAL KEMPE 175 RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (a) TYPE TIME(S) 0.6 1.0 1.6 2.0 2.6 3.0 1.6 4.0 4.6 6.0 72 28-Apr T6#2 STANDARD TERRY NORMAL STAINLESS ST. 0.6/0.9 AVERAGE 1399 1656 1891 1706 1713 1713 1721 240mm STD DEV 4.2 0.4 1.4 0.4 0.2 0.4 0.2 73 29-Apr T6#2 STANDARD TERRY FINE STAINLESS ST. 0.6/0.9 AVERAGE 1454 1663 1702 1711 1719 1731 1728 240mm STDOEV 2.1 0.8 0.4 0.4 0.4 0.4 0.3 74 30-Apr T6#2 81mm INSERT TERRY NORMAL STAINLESS ST. 0.6/0.9 AVERAGE 1292 1365 1373 1381 1386 1401 1406 240mm STD DEV 0.3 0.4 0.2 0.2 0.3 0.2 0.2 75 29-Apr T6#2 81mm INSERT TERRY FINE STAINLESS ST. 0.6/0.9 AVERAGE 1285 1339 1363 1366 1375 1380 1389 240mm STD DEV I 0.4 0.3 0.4 0.4 0.2 03 0.2 EFFECT OF SIZE AND INSERT ON T11 FEEDERS RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (a) TYPE TIME (a) 0.6 1.0 1.6 ZO 2.6 3.0 J-6 4.0 4.6 6.0 76 01-May T11 #2 STANDARD LINDBERG NORMAL STAINLESS ST. o a/o 8 AVERAGE 840 1249 1480 1596 1643 1676 1689 240mm STD DEV 3.1 1.4 0.6 0.4 0.4 0.4 0.6 77 08-May T11 #2 STANDARD LINDBERG FINE STAINLESS ST. 0.8/0.8 AVERAGE 842 1308 1631 1614 1665 1690 1701 240mm STDOEV 2.8 1.2 0.4 0.4 0.2 0.2 0.3 78 07-May T11 #2 81mm INSERT LINDBERG NORMAL STAINLESS ST. 0 8/0 B AVERAGE 833 1326 1345 1357 1362 1367 1369 240mm STD DEV 3 0.4 0.4 0.4 0.2 0.3 0.3 79 01-May T11 #2 81 mm INSERT LINDBERG FINE STAINLESS ST. 0.8/0.8 AVERAGE 771 1321 1355 1355 135S 1356 1364 240mm STD DEV | 7.4 0.3 0.3 0.2 0.2 0.2 0.5 EFFECT OF BATH SYSTEM DUST ON STANDARD AND INSERT FEEDERS RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIM: EFFECT OF STROKE DISTANCE WITH INSERT (PARKER Vs TERRY. FINE Vs NORMAL) RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (a) TYPE TIM: (a) 0.6 1.0 1.5 2.9 2-6 3.0 3.6 4.0 4.6 6.0 102 22-Jul PA #2 STANDARD PARKER FINE KEMPE 175mm * 1.6/1.9 AVERAGE 866 1559 1695 1670 1684 1729 1753 1806 1733 1812 30mm SPACER STDOEV 4.9 1.9 1 0.6 1 0.5 0.9 0.9 1 0.8 103 24-Jii PA #2 STANDARD TERRY FINE KEMPE 175mm* 1.8/1.9 AVERAGE 535 1489 1724 1654 1766 1826 1702 1886 1711 1688 30mm SPACER STD DEV 3.2 1.2 0.9 0.4 0.8 1 0.4 1.4 0.8 0.4 104 24-Jul PA #2 STANDARD TERRY NORMAL KEMPE 175mm + 1.6/1.9 AVERAGE 750 1579 1672 1691 1695 1702 1714 1715 1711 1714 30mm SPACER STDOEV 6.2 1.1 0.5 0.4 0.3 0.3 0.3 0.4 0.3 0.4 1355 1354 105 28-Jtf PA #2 53mm INSERT TERRY FINE KEMPE 175mm* 1.6/1.9 AVERAGE 468 1066 1338 1344 1344 1346 1351 1352 30mm SPACER STDOEV 2.3 3.2 0.2 0.2 0.3 0.3 0.2 0.3 0.2 0.2 106 29-Jul PA #2 53mm INSERT TERRY NORMAL KEMPE 175mm + 1.6/1.9 AVERAGE 532 1304 1391 1391 1392 1349 1339 1335 1346 1337 30mm SPACER STD DEV 1.9 1.3 0.4 0.5 0.7 0.2 0.2 0.2 02 0.2 107 04-Aug PA #2 53mm INSERT PARKER FINE KEMPE 175mm • 1.6/1.9 AVERAGE 649 1051 1346 1368 1379 1365 1377 1384 1381 1382 30mm SPACER STD DEV 2.3 2.5 1.1 0.6 0.3 0.9 0.3 0.3 0.3 0.2 1377 1377 1384 108 05-Aug PA #2 53mm INSERT PARKER NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 701 1086 1368 1367 1373 1375 1379 30mm SPACER STD DEV 2.9 1.5 0.8 0.3 0.3 0.3 0.2 0.3 0.2 0.3 109 07-Aug PA #2 STANDARD PARKER NORMAL KEMPE 175mm * 1.6/1.9 AVERAGE 775 1056 1349 1516 1543 1599 1639 1668 1688 1703 30mm SPACER STDOEV 4.2 2.2 1.9 2.1 1.5 1.6 0.9 0.8 0.5 0.4 1784 1739 110 11-Aug PA #2 STANDARD PARKER FINE KEMPE 175mm* 1.6/1.9 AVERAGE 813 1034 1233 1455 1504 1650 1598 1652 30mm SPACER STODEV 3 1.7 4.5 34 2.3 1.6 2.6 1.4 0.7 2 1463 1457 1418 111 12-Aug PA #2 81mm INSERT PARKER NORMAL STAINLESS ST. 1.8/1 9 AVERAGE 902 1439 1472 1469 1468 1456 1434 240mm STD DEV 8.9 0.4 0.2 0.3 0.2 0.3 0.4 0.4 0.3 0.4 1510 1507 1508 1507 112 13-Aug PA #2 81mm INSERT PARKER FINE STAINLESS ST 1.8/1.9 AVERAGE 742 1461 1496 1503 1504 1505 240mm STD DEV 5.9 0.8 0.4 0.2 0.3 0.4 0.4 0.3 0.5 0.5 Appendix 2 Off-site Shot Size Test Data Page 4 EFFECT OF SPRING TYPE (KEMPE 175mm + 30mm SPACER: KEMPE 210mm: SS 240m: S INSERT STROKE (81mm and 53mm); TERRY Vs PARKER RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING 3TROKE DATA DWELL TIME (a) TYPE TIME (a) 0.5 1.0 1.6 113 17-JUlg PA #2 81mm INSERT TERRY NORMAL STAINLESS ST. 1.6/1.9 AVERAGE 753 1401 1416 1416 1417 1418 1418 1418 1418 1417 240mm STDOEV 3.1 0.4 0.3 0.3 0.3 0.2 0.2 0.3 114 18-Aug PA #2 81mm INSERT TERRY FINE STAINLESS ST. 1.6/1 9 AVERAGE 687 1450 1450 1450 1449 1448 1451 1448 240mm STDOEV 5.4 0.4 0.3 0.3 03 0.3 115 17-Aug PA #2 81mm INSERT TERRY NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 878 1444 1436 1444 1442 1446 1444 1442 1448 30mm SPACER STD DEV 6.3 0.4 0.3 0.3 0.3 0.4 0.3 03 116 24-Aug PA #2 81mm INSERT TERRY FINE KEMPE 175mm * 1.8/1.9 AVERAGE 794 1446 1448 1449 1451 1449 1464 1459 1458 30mm SPACER STD DEV 3.2 0.8 0.5 0.5 0.3 0.3 03 117 26-Aug PA #2 81mm INSERT PARKER NORMAL KEMPE 175mm + 1.6/1.9 AVERAGE 867 1462 1516 1536 1486 1472 1502 1542 1487 30mm SPACER STD DEV 2.1 03 0.7 0.6 0.3 0.2 0.6 118 28-Aug PA #2 81mm INSERT PARKER FINE KEMPE 175mm* 1.6/1.9 AVERAGE 722 1521 1582 1584 1584 1584 1564 1564 1580 1581 30mm SPACER STD DEV 2.4 0.8 0.7 0.4 0.2 0.3 119 01-Sep PA (2 STANDARD PARKER NORMAL KEMPE 175mm + 1.6/1.9 AVERAGE 861 1208 1529 1673 1720 1736 1734 1744 1747 30mm SPACER STDOEV 3.3 2 2.9 0.8 0.6 0.4 0.7 0,3 0.3 02 120 03-Sep PA #2 STANDARD PARKER NORMAL KEMPE 210mm 1,6/1.9 AVERAGE 877 1221 1536 1679 1720 1734 1758 1776 1786 1783 STDOEV 1.8 1.3 1.8 1.0 1.1 0.4 05 0.6 0.2 0.4 03-Sep 121 PA #2 53mm INSERT PARKER NORMAL KEMPE 210mm 1.671.9 AVERAGE 721 1219 1336 1358 1362 1370 1368 1365 1371 1362 STD DEV 3.4 2.6 0.2 0.2 0.1 0.2 0.2 0.3 0.3 0.5 09-Sep 122 PA #2 53mm INSERT PARKER NORMAL STAINLESS ST. 1.6/1.9 AVERAGE 753 1194 1333 1369 1377 1387 1387 1391 1386 1390 220mm STDOEV 2.3 2.1 0.6 0.4 0.2 0.3 0.3 0.3 0.3 0.2 EFFECT OF CUSHION QUALITY RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (a) TYPE TIME (a) 0.6 1.0 1.6 2.0 2.8 3.0 3.6 4.0 4.5 6.0 122A 10-Sep PA #2 STANDARD PARKER NORMAL STAINLESS ST. 1.6/1.9 AVERAGE 883 1198 1560 1636 1684 1700 1742 1742 1755 1746 220mm STD OEV 2.6 2.5 1.2 2.0 0.1 1.6 0.5 0.4 0.3 0.8 123 14-Sep PA #2 STANDARD PARKER NORMAL KEMPE 175mm + 1.671.9 AVERAGE 844 1448 1531 . 1583 1844 1675 1784 1702 1700 NO CUSHION 30mm SPACER STDOEV 3.7 3.1 0.9 . 0.3 0.7 0.4 03 0.3 0.3 123 A 16-Sep PA #2 STANDARD PARKER FINE KEMPE 175mm* 1.6/1.9 AVERAGE 667 944 1282 1447 1518 1608 1836 1665 1693 1713 NO CUSHION 30mm SPACER STDOEV 3.4 3.4 3.6 2.7 2.1 3.7 2.3 1.6 1.9 OS 124 16-Sep PA #2 53mm INSERT PARKER FINE KEMPE 175mm* 1.6/1.9 AVERAGE 471 1092 1329 1341 1349 1346 . 1348 1350 1422 NO CUSHION 30mm SPACER STDOEV 2.7 6.0 0.6 0.7 0.4 1.0 - 0.3 0.5 0.8 125 21-Sep PA #2 81mm INSERT PARKER FINE KEMPE 175mm* 1.6/1.9 AVERAGE 597 1368 1426 1426 1424 1439 1442 1451 1448 1451 NO CUSHION 30mm SPACER STDOEV 2.5 1.3 0.8 0.5 0.5 0.2 0.2 0.6 0.4 0.2 126 23-Sep PA #2 STANDARD PARKER NORMAL KEMPE 175mm + 1.6/1.9 AVERAGE 780 1092 1441 1551 1631 1623 1657 1702 1712 1721 NO CUSHION 30mm SPACER STD OEV 4.8 2.9 3.9 1.6 1.3 1.9 0.8 1.2 0.3 0.2 127 29-Sep PA #2 53mm INSERT PARKER FINE KEMPE 175mm + 1.8/1.9 AVERAGE 709 1045 1285 1347 1359 1364 1366 1368 1366 1365 30mm SPACER STDOEV 2.7 3.5 2.7 0.7 0.3 0.3 0.2 0.5 0.1 0.1 128 01-Oct PA #2 STANDARD PARKER NORMAL KEMPE 175mm + 1.8/1.9 AVERAGE 809 1081 1351 1569 1604 1660 1684 1711 1713 . 30mm SPACER STD DEV 2.1 2,2 2.7 1.6 1.5 OS 0.9 0.5 0.6 . 129 OS-Oct PA #2 STANDARD PARKER FINE KEMPE 175mm * 1.6/1.9 AVERAGE 717 1055 1328 1478 1614 1658 1888 1741 1750 1746 30mm SPACER STDOEV 3.2 2.9 2.7 2.4 1.6 1.1 0.6 0.7 0.4 0.5 EX-PLANT EFFECT OF SPOOLS & SEAL AGE RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (a) TYPE TIME (a) 0.6 1.0 1.6 2.0 2.6 3.0 3.5 4.0 4.6 6.0 130 07-OM PA #2 EXPLANT PARKER NORMAL KEMPE 175mm + 1.6/1.9 AVERAGE 805 1107 1396 1628 1655 1740 1651 1681 1791 1788 1248 30mm SPACER STDOEV 2.9 3.4 3.4 1.6 3.9 1.0 0.5 1.2 0.7 1.1 131 08-Ocl PA #2 OLD SEALS PARKER NORMAL KEMPE 175mm + 1.8/1.9 AVERAGE 610 1056 1320 1349 1355 1372 1370 1388 1369 1371 53mm INSERT 1248 30mm SPACER STD DEV 2.5 2.4 2.1 1.3 0.7 0.3 0.3 0.3 0.4 0.3 132 10-Ocl PA #2 NEW SEALS PARKER NORMAL KEMPE 175mm + 1.6/1.9 AVERAGE 872 1146 1394 1509 1565 1589 1635 1861 1704 1699 STANDARD 1248 30mm SPACER STDOEV 3.1 3.0 2.0 1.3 1.6 1.2 0.6 0.3 2.6 0.3 133 21-Oct PA #2 NEW SEALS PARKER NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 696 1313 1419 1415 1392 1497 1419 1386 1415 1413 53mm INSERT 124B 30mm SPACER STD DEV 2.1 1.5 0.5 0.5 0.3 0.6 0.6 0.8 0.2 0.2 ATLAS COPCO RUN DATE FEEDER DESIGN CYLINDER ALUMNA SPRING STROKE DATA DWELL TIME (a) TYPE TIME (a) 0.6 1.0 1.5 2.0 2.5 3.0 3.6 4.0 4.6 6.0 142 27-Oct PA #2 STANDARD 9001 NORMAL KEMPE 175mm + 1.6/1.9 AVERAGE 857 1142 1511 1602 1660 1703 1730 1758 1771 1776 30mm SPACER STDOEV 3.6 2.0 1.3 1.8 1.5 1.0 0.7 0.7 0.5 0.7 143 27-Ocl PA #2 53mm INSERT 9001 NORMAL KEMPE 175mm* 1.6/1.9 AVERAGE 763 1105 1330 1359 1371 1372 1376 1388 1377 1384 30mm SPACER STD DEV 2.4 2.4 1.3 0.5 0.8 0.3 0.4 0.5 0.7 0.7 144 28-Oct PA #2 53mm INSERT 9001 FINE KEMPE 175mm* 1.6/1.9 AVERAGE 701 1109 1361 1356 1392 1418 1423 1419 1419 30mm SPACER STD DEV 2.8 1.8 1.8 1.1 1.1 0.5 0.5 0.5 0.4 145 28-0« PA #2 STANDARD 9001 FINE KEMPE 175mm* 1.6/1.9 AVERAGE 878 1181 1490 1638 1770 1794 1819 1861 1868 1881 30mm SPACER STD DEV 3.7 3 3.8 1.9 0.6 1.3 0.6 0.6 0.8 0.5 Notes: (I) All shot size mass units are g (ii) Average and standard deviations are based on 25 consequitive shots of alumina. (iii) Stroke time is "downstroke time/upstroke time" in s. APPENDIX 3 ALUMINA SIZE AND DENSITY Appendix 3 Alumina Size and Density Pagel (a) ALUMINA SIZING (a) NORMAL ALUMINA RUNS (JAN - SEP.92) (b) FINE ALUMINA RUNS (JAN -AUG,92) RUN +100# +325# D10 D50 090 <21 FFT RUN +100# +325# D10 D50 D90 <21 FFT MICRON MICRON 7 8.0 96.8 52.4 114.1 172.1 5.0 1 8.0 96.8 18.1 105.8 167.1 10.0 10.0 96.4 47.9 112.7 170.6 9 8.0 5.6 3 7.3 96.0 16.0 104.3 165.6 12.9 10 7.2 96.8 60.5 115.8 173.0 5.2 4 7.2 94.4 33.5 109.4 171.6 7.3 11 7.6 96.0 61.1 115.3 178.3 4.5 5 6.8 96.0 36.2 109.6 173.1 7.4 12 6.4 96.4 56.0 114.2 169.9 5.3 6 8.0 96.8 34.2 110.1 170.4 7.2 22 8.8 93.6 56.5 112.3 171.3 4.2 3.8 19 6.8 96.0 60.7 116.7 181.3 6.0 5.2 23 8.8 96.4 68.9 117.1 179.3 3.8 20 7.2 96.0 38.5 111.2 172.4 5.7 25B 6.8 96.4 65.4 114.2 168.3 3.1 3.5 25A 8.8 96.4 57.1 112.3 168.4 4.4 4.5 26 8.4 97.2 63.2 107.1 158.3 1.5 3.6 29 8.0 98.0 57.1 110.5 164.9 4.1 4.3 27 9.6 98.0 70.1 112.6 170.4 3.9 3.4 37 9.6 98.4 59.0 110.0 164.5 3.7 3.6 28 8.4 99.2 69.5 112.3 169.8 1.2 3.5 38 9.2 98.0 57.2 109.9 165.0 4.1 4.1 30 8.3 96.9 74.7 119.8 173.4 1.7 3.4 39 8.8 98.0 57.9 110.2 166.0 4.1 4.2 31 8.2 97.5 71.4 113.9 171.8 3.2 47 8.8 97.2 63.9 110.5 163.1 2.7 3.5 32 7.9 97.4 77.3 119.1 177.2 3.3 49 8.4 97.2 69.2 112.0 170.0 1.1 3.4 33 8.2 97.5 71.4 113.9 171.8 3.2 52 9.6 96.4 62.6 113.2 170.9 3.7 3.9 34 8.9 97.4 95.5 118.0 177.8 3.3 53 9.2 97.2 65.5 114.2 167.1 3.1 3.8 35 9.6 98.0 70.1 112.6 170.4 3.9 3.4 55 7.6 97.2 71.9 113.5 171.2 3.8 3.3 36 10.0 97.6 1.2 3.4 56 7.6 98.0 74.9 131.7 373.4 3.7 3.3 40 8.4 97.1 75.7 117.8 177.3 0.8 3.3 59 10.8 98.4 74.4 116.7 169.0 3.8 3.3 42 7.6 98.8 69.6 110.1 176.6 3.5 3.2 61 8.4 97.6 72.3 115.2 167.1 3.7 5.4 44 8.8 99.2 73.4 113.1 163.9 3.5 3.2 62B 7.6 98.0 73.3 114.9 166.2 3.8 3.3 46 8.0 98.8 69.9 110.4 167.0 2.8 3.2 67 7.6 98.4 71.0 112.3 163.2 3.7 3.3 50 8.8 98.8 76.5 113.7 163.1 3.3 3.2 70 10.6 98.3 72.5 113.7 173.1 3.4 3.3 54 11.2 98.8 72.7 113.4 170.8 3.4 3.3 74 9.6 98.4 75.4 116.5 176.1 3.3 3.3 57 7.6 98.0 74.6 115.7 166.5 2.6 3.3 77 9.1 98.3 75.1 115.6 173.5 3.4 3.2 58 8.0 98.4 78.8 120.8 174.9 2.7 3.3 79 10.2 99.1 75.1 115.7 170.1 3.2 3.2 60 8.4 98.8 3.5 3.2 AVE 8.5 97.3 58.6 112.9 177.1 3.8 4.8 62A 10.8 98.8 76.7 116.4 169.4 3.4 3.3 S.D. 1.1 1.1 17.6 4.9 39.5 1.8 2.3 63 8.8 98.8 73.6 113.8 170.5 3.5 3.2 102 8.1 97.5 40.7 107.0 170.5 7.3 7.2 64 8.8 98.4 70.6 111.2 161.9 2.7 3.2 103 9.4 95.0 39.7 105.2 169.4 7.3 6.2 66 8.8 99.2 3.4 3.2 105 6.5 95.4 52.3 107.2 166.5 5.2 4.7 71 10.5 97.9 76.1 116.9 175.6 3.4 3.3 107 6.3 95.6 57.7 109.5 174.1 4.1 4.2 72 10.5 99.2 82.0 120.4 172.5 3.4 3.2 110 7.3 96.4 53.0 105.1 163.6 4.4 4.7 76 10.2 98.9 77.0 117.7 170.8 3.4 3.2 112 7.9 96.2 57.5 106.0 164.3 4.1 4.7 78 9.6 98.3 74.5 115.0 172.4 3.4 3.1 116 6.7 96.4 54.3 104.7 165.4 4.0 4.7 AVE 8.7 97.8 70.4 114.7 171.2 2.8 3.6 118 7.4 97.4 54.2 104.5 160.7 4.1 3.4 S.D. 1.1 1.2 9.1 3.1 4.8 1.1 0.7 AVE 7.5 96.2 51.2 106.2 166.8 5.1 5.0 101 5.0 96.0 50.7 107.5 174.2 5.7 4.9 S.D. 1.0 0.8 6.6 1.6 4.0 1.3 1.1 104 7.5 97.0 53.3 105.5 163.9 5.0 4.4 106 6.5 98.0 59.9 107.6 169.6 3.4 3.9 108 6.2 97.4 70.0 112.2 170.1 1.7 3.7 109 7.0 97.6 56.0 107.5 154.5 1.6 3.6 111 7.6 97.1 66.3 109.4 166.8 1.3 3.5 (c) UNREACTED ALUMINA RUNS (JAN.92) 113 6.6 97.1 70.3 110.9 165.1 1.0 3.5 115 7.4 97.4 65.5 106.0 162.5 1.2 3.4 117 8.0 97.9 66.4 106.2 166.5 1.2 3.7 RUN + 100# +325# D10 D50 D90 <21 FFT 119 7.5 98.0 65.4 104.8 161.3 1.1 3.5 MICRON 120 8.0 97.6 59.4 109.4 166.7 0.6 3.4 13 9.4 94.2 36.5 113.6 170.5 6.1 121 7.2 97.3 72.5 113.1 171.6 0.7 3.4 15 7.8 93.9 38.6 113.6 170.2 6.5 122 8.0 97.7 67.7 106.5 162.2 0.6 3.3 16 6.8 94.8 46.4 113.3 174.8 5.4 122A 8.3 98.1 70.8 106.6 166.4 0.5 3.4 17 7.6 96.8 30.3 106.9 172.9 5.1 123 7.0 98.5 67.2 105.2 161.4 0.4 3.3 18 7.6 94.8 54.1 115.4 174.1 5.5 AVE 7.2 97.5 64.1 107.9 165.5 1.7 3.7 AVE 7.8 94.9 41.2 112.6 172.5 5.7 S.D. | 0.8 0.6 6.5 2.5 4.7 1.6 0.4 S.D. 0.9 1.0 8.3 2.9 1.9 0.5 Notes: (i) +100* and +325# refer to Tyler mesh. Values are cumulative percent oversize. (ii) D10, D50 and D90 refer to the micron size above of which 10%, 50% and 90% of the sample is gre (iii) "FFT" is funnel flow time ie. the time in minutes that 10Og of alumina flows from a standard funnel. The longer the time, the finer the size. This is a standard analysis within the alumina industry. There is almost a linear relationship between sub 20 micron sizing and FFT. Appendix 3 Alumina Size and Density Page 2 (b) ALUMINA DENSITY COMPARISON (RUNS 1-18) ALUMINA LOOSE PACKED RUN TYPE DENSITY DENSITY NORMAL 1030 1163 22 1026 1153 23 1020 1200 7 1020 1220 9 1033 1180 10 1030 1181 11 1034 1185 12 AVERAGE 1028 1183 STD DEV 5 21 ALUMINA LOOSE PACKED RUN TYPE DENSITY DENSITY FINE 1030 1270 1 1040 1270 3 1058 1186 4 1055 1185 5 1063 1203 6 1034 1210 19 1030 1182 20 AVERAGE 1044 1215 STD DEV 13 36 ALUMINA LOOSE PACKED RUN TYPE DENSITY DENSITY UNREACTED 990 1200 13 980 1210 15 1008 1157 16 1015 1145 17 1008 1146 18 AVERAGE 1000 1172 STD DEV 13 28 Note: Density measured as kg/m3 APPENDIX 4 BOB SEYMOUR NOTE ON FEEDER RECOMMENDATIONS Appendix 4 Bob Seymour Memo on Feeder Recommendations Pagel R. E. Seymour Smelting Managers Primary Metals Knoxville 12 1992 April 16 Alumina Feeder Improvement A "main-ring event" at the Feed Alumina Network meeting in Wanatchee April 9-10 was art update by Jim Kissane (Portland) on alumina feeders. Earlier work was contained in the report Pot Alumina Feeder Design Review. J. P. Kissane. August 1991. Jim did an excellent job of informing Feed Alumina contacts using a standup presentation, a notebook to document findings, and a videotape. The best companies in the world are able to operate alumina feeders for 5-7 years (the life of the cathode) without replacement. Alcoa plants range from 6 to 36 months. We are definitely not a benchmark. It appears that Karmoy (Norsk Hydro), who holds the benchmark for anode effect frequency (0.03 AE/PD), also holds the benchmark for feeder life (about 7 years). I wonder if that is coincidence? As i see it, Alcoa has managed feeders even worse than we have managed pot life. We know the average life to failure, but little if anything about failure age distribution or root causes. Not unlike potlining, our focus has been on improving change out procedures rather than systematically eliminating causes of failure. Attached are recommendations from the work done at Portland that can be acted upon. Most recommendations apply at all locations, although there may be exceptions. Here is an opportunity to rapidly transfer technology important to a Quantum Leap Milestone. One common theme in the recommendations is targets. I suspect that few locations have ever measured Feeder Downstroke Time or Feeder Dwell Time and that even fewer locations have specification nominals and tolerances. It would seem to be a natural for pot operators to check these characteristics when they investigate multiple anode effects. Sampling and control charting or regular auditing also seem to be in order. Appendix 4 Bob Seymour Memo on Feeder Recommendations Page 2 Feeder Improvement Recommendations 1. Establish a system to track alumina feeder maintenance by failure type, pot, and component and a performance measurement system using Pareto and Weibull analyses. Portland has developed a tagging and computer system that is to be adopted by Point Henry. There is probably an advantage in adopting the Portland classification of failures, since we could then compare performance. 2. Establish a system to require suppliers of critical feeder components - pluggers, seals, springs, etc - to provide evidence of that product specifications are being met Are their manufacturing processes in control and capable? 3. Measure distribution of feeder downstroke time (time from solenoid energized to cylinder full down) for each feeder system. Establish downstroke time targets for each system and reduce variability around the target. A suggested target is: Downstroke Time - 1.5 seconds Tuning will be required to assure that the crust is broken reliably. Each combination of pot type, cylinder type, and valving is a unique system. Causes of variability can include blocked mufflers, piston seal leaks, solenoids, etc. Avoid tuning the system to take care of special cases - pots with leaky seals or plugged mufflers. High cylinder velocity increases wear by a cubed factor. 4. Survey blocked (mounded, dry packed) feeders and analyze the causes several ways including potroom location such as section or room, work activity, time of day, and anything else you can think of. 5. Reduce feeder dwell (time that plugger is in full down position) to minimize plugger wear. A suggested target is: Dwell Time - Downstroke time + ( 1.2 to 1,5 seconds). 6. Test dual dwell times - sv^ry third feeder cycle has a long dwell dwell time to break the crust. The first two cycles are short dwell times - just enough to let the alumina get dumped. Not all systems may have the hardware to implement dual dwell times. Appendix 4 Bob Seymour Memo on Feeder Recommendations Page 3 7. Establish measurable criteria for new/rebuilt feeder performance and establish a system of quality checks to assure that criteria are met. Spring length, air cushion adjustment, plugger diameter, rod seal leakage, piston seal leakage and mounting bolt tightness (loose bolts cause arcing) are important criteria. 8. Install "spool inserts" in feeders on a change out basis to make feeder shot weights more robust to changes in alumina flowability. Portland will recommend a design for each plant 9. Change specifications for pot pluggers to the following: Chemical Composition Element C Si Mn Cr Mo S P Cu Nominal 3.1 1.7 0.7 1.4 0.5 0.0 0.0 0.16 USL 3.2 1.4 0.5 1.2 0.4 0.1 0.1 0.18 LSL 3.0 1.2 0.8 1.5 0.6 0.0 0.0 0.20 Hardness BHN(300O) Nominal 350 USL 400 LSL 300 NOTE: The new specification will produce a much harder iron that cannot be machined. The holes will have to be cast in place, or a steel body will have to be cast into the plugger. Shape Flat Bottom 10. Survey pots for piston seal leakage, which contributes to high air consumption and low air pressure. 11. Implement location specific recommendations relating to piston seals, washers, springs, and other cylinder components. These can be found in the notebok supplied to your location by Jim Kissane. Appendix 4 Bob Seymour Memo on Feeder Recommendations Page 4 A reasonable goal for feeder life should be to equal the goal for potlife. Improved feeder life will support quantum leap improvements In anode effects, power consumption, and safety. In addition, longer feeder life offers the opportunity to substantially reduce maintenance costs. USMS plants should review the attached recommendations, the Portland videotape and the summary from Jim Kissane's notebook. Be prepared to discuss your ideas and plans for these recommendations during the May round of SSLT visits. For each aspect of feeders - purchased components, rebuilt superstructures, operation, reliability - accountability should be clearly defined. Some questions to oonsider are: Who Is accountable? What systems make accountability operational? What data is monitored? R. E. Seymour distribution: ^melting Managers L.E. Tate - Badln V.C. Adorno - Massena W.J. Drake - Rockdale CJ. Siizewski - Tennessee D.W. Willett - Warrick D.J Carney - Wenatchee G.A Turnbow - Sao Luis D.G. Judd - Portland L.B, Davey - Point Henry copies: G.J. Pizzey - Melbourne F.G- Tigre - Sao Paulo B.C. Raw© - Knoxville 12 R.R. Taylor - Knoxville 12 K.A. Isakson - Knoxvilte 10 J.L. Roddy - Knoxville 10 C.A. St. Clair - Knoxville 10 APPENDIX 5 DICK TAYLOR NOTE ON RESULTS OF FEEDER RESEARCH Appendix 5 Dick Taylor Memo on Results of Feeder Research Pagel ALUMINUM COMPANY OF AMERICA 12W RfVgRV^W TOWER 900 SOUTH GAY m~m£~ KN0XVSJJ;. T£H?C0SE£ 37902~984!J a 1994 July 25 ALCOA HE: JAMES P. KISSANE - PORTLAND The efficiency 0* the eieeueiytie process of converting aluminium oxide (alumina) to aluminium is primarily dependant on the control of alumina thai is dissolved in the cryolite bath {electrolyte) The ideal percent of alumina in solution is approximately 2.0 to 2,5 percent in most cells- When the alumina in solution increases significantiy above the ideal level, the process becomes unstable and unproductive for a long period of time, When the alumina in solution decreases below the idea percent, the resistance of the ceil dramatically increases to a level requiring immediate corrective action. This latter condition is named an 'anode effect". Fortunately, an anode effect can be readily corrected. Historically, the electrolytic process has been controlled by reducing the percent alumina in solution below ideal forcing an anode effect several times per day. During an anode effect when the cell resistance increases and the pot voltage exceeds approximately 8 volts tetrafiuromethane and hexaflluooethane gases are generated. These gases are not generated during normal operation of the celL Studies by atmospheric scientists indicate that these gases contribute to global warming. A program was established at the Alcoa international smelting locations to reduce the number of anode effects per operating cell day •»% -to 1990- An objective suggested by the United States EPA was a 30% to 60% reduction of anode effects per cell day by the year SQO0. Improved operating efficiency was a secondary objective of this program. Alumina is fed into the cett by an alumina feeder which is controlled by a computer based on cell resistance, The functionalfy of the alumina feeder is a key enactor to the control of the percent alumina in solution and, consequently, a key enablerto th e reduction of anode effects. Each of the nine international locations Initiated a program to improve the functtonaJty of the alumina feeder. James P. Kissane participated in this activity. Over time and on his own Initiative, Jim Kissane conducted an in-depth research project on the functionality of the feeder and also on the computer control logic and, to some extent, the related ceil chemistry. He assisted in establishing an alumina feeder information network between the international locations and travelled extensively conducting investigations and seminars. Jim is recognized as the leader of our programs related to alumina feeders. Appendix 5 Dick Taylor Memo on Results of Feeder Research Page 2 James P. Kissane Page Two 1994 July 25 The improvement in the design, control and functionality of the alumina feeder is a success story and we recognize Jim Kissane as the primary leader of this program. More importantly has been his contributions to the understanding and philosophy of cell operation relative to alumina feeding and the feeder control functions of the computer control system. Jim Kissane's research activities have been a key enabler to achieving the reduction of anode effects per cell day in the international locations and, consequently, the significant reduction of perfluorocarbon emissions. A graph indicating our experience in minimizing anode effects per cell day is enclosed for your reference. Sincerely, /^£&*^*JSI A~* yz Richard R. Taylor Director of Technology Primary Metals Division ANODE EFFECTS PER POT DAY INTERNATIONAL PREBAKES 1.5 1 Q 0.5 BASE JFMAMJJASONDJFMAMJJASONDJFMAMJJASONDJFMAMJ 1991 1992 1993 1984 1995 \SHARE\BISHOPVAlFUKWHUNTAE-PD APPENDIX 6 A2 U.S.A. PATENT APPENDIX 6 A2 U.S.A. PATENT PAGEI The Commissioner of Patents and Trademarks Has received an application for a patent for a new and useful invention. The title and description of the invention are en closed. The requirements of law have been complied with, and it has been de termined that a patent on the invention shall be granted under the law. Therefore, this United States Patent Grants to the person or persons having title to this patent the right to exclude others from making, using or selling the invention throughout the United States of America for the term of seventeen years from the date of this patent, sub ject to the payment of maintenance fees as provided by law. Commwionet of Patents aad Trademarks Amu APPENDIX 6 A2 U.S.A. PATENT PAGE 2 US005324408A United States Patent [19] [it] Patent Number. 5,324,408 Kissane [45] Date of Patent: Jun. 28, 1994 [54] APPARATUS FOR CONTROLLED SUPPLY OF ALUMINA FOREIGN PATENT DOCUMENTS [75] Inventor: James P. Kissane, Portland, 31274/71 7/1971 Australia . Australia 66481/74 3/1974 Australia . 73198/81 7/1981 Australia [73] Assignee: Portland Smelter Services Pry. Ltd., 14784/88 10/1988 Australia Victoria, Australia OTHER PUBLICATIONS [21] Appl.No.: 30,167 Patent Abstract, Metallurgy, p. 18, Aliunimum electro [221 PCT Filed: Apr. 29,1991 lyser 83-806124/44. (Dec 1982). Patent Abstract, Metallurgy, p. 23, Altuninjrum produc [86] PCT No.: PCT/AU91/00169 ing electrolyser .... 90-374562/50 (Apr. 1990). § 371 Date: Apr. 2, 1993 Primary Examiner—Donald R. Valentine § 102(e) Date: Apr. 2,1993 Attorney, Agent, or Firm—Birch. Stewart, Rolasch & Birch [87] PCT Pub. No.: WO92/06230 [57] ABSTRACT PCT Pub. Date: Apr. 16, 1992 Apparatus for the controlled supply of alumina to an [30] Foreign Application Priority Data electrolysis tank having an electrolyte crust breaking Oct. 5, 1990 [AU] Australia PK2658 plunger (2) includes a supply chamber (19) connected to 5 the entry port (16) of a dose holder (10). Alumina leav [51] Int. a. C25C 3/14 ing an exit port (15) of the dose holder (10) passes via an [52] U.S. CI 204/245 inclined wall (9) to a delivery chute (21) which directs [58] Field of Search 204/67, 243 R-247 the alumina to a hole formed in the crust by the plunger. [56] References Cited Valve means (14) movable with the plunger (2) controls the opening of the dose holder entry and exit ports (16, U.S. PATENT DOCUMENTS 15), closing one port as it opens the other, and allowing 3,371,026 2/196S Kiley et al 204/245 alumina to flow through the delivery chute (21) as the 3,901.787 8/1975 Niizeki et al 204/245 plunger (2) is retracted from the crust. The plunger 4.049.529 9/1977 Golla 204/245 X movement required to control the valve means is such 4,321,115 3/1982 Rebmann et al 204/245 X that alumina can be fed into the tank substantially con 4,328,085 5/1982 Friedl et al _. 204/245 tinuously without meeting interference from the 4,435,255 3/1984 Casdas 204/245 X plunger. 4,437,964 3/1984 Gerphagnon et al 204/245 4,525.105 6/1985 Jaggi 204/245 X 5 Claims, 1 Drawing Sheet 5,045,168 9/1991 Dalen et ai 204/245 APPENDIX 6 A2 U.S.A. PATENT PAGE 3 U.S. Patent June 28, 1994 5,324,408 APPENDIX 6 A2 U.S.A. PATENT PAGE 4 5,324,408 1 2 crust formed on the surface of molten electrolyte, the APPARATUS FOR CONTROLLED SUPPLY OF crust breaking mechanism including a plunger with a ALUMINA cutting edge mounted on a reciprocable plunger shaft, and an alumina storage container adapted to release BACKGROUND OF THE INVENTION 5 alumina as required for entry into the electrolyte I. Field of the Invention through the hole in the crust, characterized in that the This invention relates to an apparatus for the con- storage container feeds alumina through an alumina trolled supply of alumina or other solid materials to an supply passage and an entry port into a supply chamber electrolytic tank in which the alumina is converted to defined between an inner wall of the feeder assembly aluminium. 10 an^ an outer supply chamber wall; a supply chamber 2. Description of Related Art exit port controlled by a valve means connects the sup- In the electrolysis of alumina, solid alumina is dis- ply chamber to a dose holder having an inner wall solved in a tank or pot containing a molten electrolyte mounted around and concentrically with the plunger such as cryolite and it is desirable to maintain the alu- shaft; the inner wall is urged downwardly towards the mina concentration in the electrolyte within a predeter- " head of the plunger; an entry port in the dose holder is mined range. In current practice for the electrolysis of immediately adjacent to the supply chamber exit port so alumina, the alumina is fed in successive doses of prede- that when the valve means opens the supply chamber termined size into one or more holes which are made in exit port, it simultaneously opens the dose holder entry the electrolyte crust so that the alumina can be admitted port and alumina in the supply chamber is able to flow when required. As the electrolysis of the alumina pro- 20 directly to the dose holder, the valve means is opera- ceeds continuously, it would be desirable if the alumina pvely associated with the inner wall so as to move in consumed in the electrolysis process could be continu- response to the movement of the inner wall between a ously replaced so as to maintain the optimum alumina {j^ position in which the dose holder is closed to the concentration in the electrolyte. However, the optimum supply chamber and a second position in which the dose operating conditions are such that the electrolyte crust 15 holder is opened to the supply chamber, the valve continuously reforms on the surface of the electrolyte means being open in itsfirst position to a flow passage making it difficult to continuously supply alumina to the defined between the inner wall and the valve means and molten electrolyte beneath the crust. For this reason, m {ts ^^4 position closing off the dose holder from known alumina feeding procedures involve the use of a the fl 5,324,,40 8 3 4 finely divided alumina. Other storage containers may be exit port leading to the flow passage or an entry port associated with the feeder assembly for other additives leading to a supply chamber. The nature of the port in to the electrolysis tank such as aluminium fluoride, the dose holder is controlled by the movement of the calcium fluoride, crushed bath, soda ash, or cryolite. valve means so that when either port is fully closed, the The other storage containers may be adapted to feed 5 other is fully open. their contents into the tank in a similar manner to that Preferably the lower end wall is substantially down described below for the alumina. wardly and inwardly inclined at an angle greater than A supply chamber provided between the storage the angle of respose of the alumina powder which is to container and a dose holder includes a preferably sub be fed through the dose holder. This inclination of the stantially cylindrical inner wall mounted around and 10 lower end wall ensures that all the alumina powder concentrically with the plunger shaft. The inner wall is (other than that held in the annular seat) will flow from urged downwardly towards the head of the plunger, the dose holder when the exit port is open. preferably by spring pressure exerted between a radially The inclination of the upper end wall is substantially outwardly extending flange on the inner wall and a downwardly and outwardly. The upper end wall is feeder assembly outer wall which is also mounted con- 15 preferably also inclined at an angle greater than the centricaUy with the plunger shaft The feeder assembly angle of repose of the alumina powder which is to be outer wall may include a radially extending flange more fed through the dose holder. This inclination of the remote from the plunger head than the flange on the upper wall ensures that the dose chamber will be filled inner wall so that a coil spring mounted between the with alumina, thus providing the desired accurately respective inner wall and outer wall flanges can exert 20 reproducible dosage. the desired pressure urging the inner wall downwardly The annular seat in the lower end wall not only pro until its downward movement is terminated. The spring vides a means of sealing the exit port of the dose holder. is mounted in the upper portion of the supply chamber It also provides a stop to terminate the downward so that alumina in the supply chamber will not interfere travel of the valve means and the associated inner wall with the spring operation. 25 which occurs when the plunger shaft is lowered in The supply chamber is defined between the inner response to the downward urging of the spring or other wall of the feeder assembly and a preferably substan pressure exerting means. The valve means is held in the tially cylindrical outer supply chamber wall. The sup lower end wall seat by the downward pressure while ply chamber includes an entry port connected to an the plunger shaft may be driven further downwardly if alumina supply passage below the inner wall flange and 30 the crust is to be broken. an exit port controlled by a valve means. The capacity When the plunger shaft is raised, means consisting of of the supply chamber is preferably at least that of the the plunger head itself, or the preferred striker means, dose holder. The inner wall at the supply chamber is meets the lower edge of the inner wall and raises it and preferably supplemented by a substantially down the associated valve means to close the entry port and wardly and outwardly directed supply chamber side 35 open the exit port of the dose holder. The upward wall which terminates at its lower edge by the supply movement of the inner wall is terminated when the chamber exit port. Preferably the supply chamber side upper end edge of the valve means seats within the wall is inclined at an angle greater than the angle of annular seat in the upper wall of the dose holder. repose of the alumina which is to pass through the sup The feeder assembly further includes an inclined wall ply chamber. This ensures that the alumina will flow 40 connected adjacent to the lower end of the inner wall. freely through the chamber. The inclined wall is preferably of substantially frusto- The supply chamber exit port is immediately adjacent conical form and terminates at its lower, free edge at or to an entry port in the dose holder so that when the within the entry portion of a delivery chute. valve means opens the exit port of the supply chamber, The delivery chute is adapted to be mounted below it simultaneously opens the entry port to the dose 45 the feeder assembly and is adapted to provide a funnel- holder, and alumina in the supply chamber is able to like action to direct alumina which leaves the dose flow directly to the dose holder. holder to one or more outlets which terminate in use The valve means is operatively associated with the above the hole in the electrolyte crust. The delivery inner wall so as to move in response to the movement of chute preferably directs all the alumina leaving the the inner wall between a first position in which the dose SO lowar edge of the inclined wall at the base of the inner holder is closed to the supply chamber, and a second wall, towards one or more delivery outlets 5. position in which the dose holder is opened to the sup To assist a further understanding of the invention, ply chamber. In its first position, the valve means is reference is now made to the accompanying drawing open to a flow passage defined between the inner wall which illustrates one preferred embodiment of the pres and the valve means. In its second position, the valve 55 ent invention. It is to be appreciated that this embodi means closes off the dose holder from the flow passage. ment is given by way of illustration only and that the The valve means is preferably substantially cylindrical invention is not to be limited by this illustration. and is connected to the inner wall between itsfree end BRIEF DESCRIPTION OF THE DRAWING edges. Each of the respective free end edges of the FIG. 1 is a diagrammatic sectional view of the pres cylindrical valve means is adapted to seat in an annular 60 ent invention. seat defined at the opposite ends of the dose holder. DETAILED DESCRIPTION OF PREFERRED The dose holder is a chamber defined by an outer EMBODIMENT wall which is preferably substantially cylindrical and The drawing shows, somewhat diagrammatically, has two radially inwardly directed end walls in which one half only of a sectional view of a preferred form of the respective annular seats are defined, and a radially 65 feeder assembly. Plunger shaft 1 is connected to plunger inward movable wall formed by the valve means. De 2, and shoulder 3, which is at the junction of plunger pending on the position of the valve means, the dose holder will always include an open port constituting an APPENDIX 6 A2 U.S.A. PATENT PAGE 6 5,324,408 5 6 shaft 1 and plunger 2, abuts striker means 23 on inner L A feeder assembly for an alumina electrolysis tank wall 4 in the position shown. Inner wall 4 is urged including a crust breaking mechanism operable to break downwardly by spring 5 which is held between flange 6 a hole in crust formed on the surface of molten electro- on inner waD 4 andflange 7 on outer wall 8. Inclined lyte, the crust breaking mechanism including a plunger wall 9 at the lower end of inner wall 4 is connected 5 with a cutting edge mounted on a reciprocable plunger adjacent to the junction of inner wall 4 and the striker shaft, and an alumina storage container adapted to re- ^^ „ lease alumina as required for entry into the electrolyte The chamber forming dose holder 10 is defined be- through the hole in the crust, characterized in that the tween side wall 11 and end walls 12 and 13, together storage container 8 adapted to feed alumina Ora-j|h»o wnTvalve means 14 which comprises the moveable "> d™?"**JW "* «Wpo«%°* £«£ wall connected to inner wall 4. In the drawing, valve chamber defined between an inner wjdl of the feeder open, while entry port 16 is closed. Annular seats 17 and ^ dose holder, the inner wall is 18 for me respectiveend edges o^ta: meanswall 14 ^^J^S^ «J2S*tf y with the plunger are formed m the respecOve end walls 12 and 13. ^ ^ ^ waU a Supply chamber 19 isfilled generally below the levd ^,^wardl toward$ ^j,^ of ^ pi^ger; an offlange 6 by alumma entering as mdicated by arrow 20 • fa ^ dQsc hoJ 5,324,408 8 than the angle of repose of the alumina powder which is annular valve seats formed in the upper and lower end walls of the dose bolder. to be fed through the feeder assembly. 5 The feeder assembly as claimed in claim 1, wherein 4. The feeder assembly as claimed in claim 1, wherein the supply chamber is formed with an inclined inner the valve means is substantially cylindrical, is connected wall which terminates at its lower edge by the supply to the inner wall between its free end edges, and seats in chamber exit port. U.S. Patent June 28, 1994 5,324,408 APPENDIX 7 PULSE CHUTE PATENT APPLICATION APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE (Th« foilowioi 11 io Dt ruled (a oy me receiving Office i INTEaWATlONAL INTERNATIONAL APPLICATION APPLICATION Ne: UNDER THE INTERNATIONAL PATENT COOPERATION TREATY RUNG DATE: REQUEST THE USDE15IG1WD BEQUESTS THAT THE PlESEnT (Stamp/ INTERNATIONAL APPLICATION BE PIOCESSED _Nameof receiving Office and 'PCT International Application" ACCORDING TO THE PATENT C0OPEIAT1ON TBEATY Applicants or Agent's File Reference (indicated by applicant if desiredl B.i No. I TITLE OF INVENTION METHOD AND APPARATUS FOR CONTINUOUS SUPPLY OF ALUMINA BoiNo.II APPLICANT (WHETHER. OR NOT ALSO INVENTOR); DESIGNATED STATES FOR WHICH WF/c.HF/iTrc The person identified in this box is (check one only)- I { applicant and inventor* Name aaaf address:** PORTLAND SMELTER SERVICES PTY. LTD. 535 BOURKE STREET, MELBOURNE, VICTORIA 3000, AUSTRALIA Telephone number: Telegraphic address: Teleprinter address (including area code) Country of nationality: AUSTRALIA Country of residence-*" The person identified m this box is applicant for the purposes of (check one only): i tall designated States (Ti''1 designated States except j [the United States i 'the Slates indicated 1—I L£_ithe United Sutes of America i——lof Arnenca only I Jiiwn the 'Supplementa"Supplemental Boi' Bar No. 01 FURTHER APPLICANTS. IF ANY: (FURTHER) INVENTORS. IF ANY; DESIGNATED STATES FOR WHICH THEY ARE APPLICANTS (IF APPLIC AB LE). A icparalc sub-box has to be filled in in respect of each person i includes wnere applicable, a legal entity), if the following two sub-boxes are insufficient, continue in the "Supplemental Box." (giving there for each addi tional person the sane indications as (hose requested in the following two sub-boxes) or by using a "continuation sheet." The person identified in this tub-box i> (check one only) |Xj applicant and inventor* f~J applicant only mvemoronly" Name anal address:** KISSANE, Janes Patrick 3 MOFFATT COURT, PORTLAND, VICTORIA 3305, AUSTRALIA If the person identified in this sub-bjf is af pliant (or applicant and manor j. indicate also: Country of nationality: AUSTRALIA Country of residence:**" and whether that person is applicant tot the purposes of I check one only): Gall deaienatact Statsta I l»" designated States except [~7]the United Stales j Ithe States indicated * I I the United Stales of America Lilof America only I lin the "Supplemental Box" The person identified in this sub-box i»< check one only I ) [ applicant and inventor" 1 applicant only , ' mvemoronly* Name ana addreu:*" If the person identified in this sub-oox is applicant tor applicant an4 micntor), indicate also: Country of nationality Country of residence:*** and whether thai person is applicant tot the purposes of (check one only): .u dnisBsatid suu. 1 1*" »«*Mjneied Susies eaeept I I the United Stales I I the Slates indicated • all designated Stales | [^ •jrJ2*djtiltl ^ AmJ!nca 1 lof America only I—I in the "Supplemental Box If the perxaat mdr utrl at atypiicaeu end inventor" or as -inventor only" is not aa lanwauerfbr the purposes of all the designated States. free the necessary wdfcauoaa in UM -Svpsteeneatal box.* Indicate the name ofanattirel person by|hriB«hii/herfamilyiiamtni*foUowe^ . ijs full official designation la the address, include both the postal coda (if any) and the country (name). If residence if not indicated, il will be assumed that the country of residence it the sane as the country indicated in the address. FonarXT/RO/101 Sheet number ff W COMMON : CERTAIN CA^< J )°» "* ""ENTATIVE(IFANY); ADDRESS FOR NOTIFICATIONS IN ^ha^e^eX^ *•" *»""«•<' a. agent or common representative » a:, Name ana address, including postal code and country: If the space below is used instead for an PHILLIPS ORMONDE & FITZPATRICK address for notifications, mark here 367 COLLINS STREET, MELBOURNE, VICTORIA 3000, AUSTRALIA Telephone number. (°3) 614 1944 Telegraphic Teleprinter • including area code) address: address. OR TREI™0!*!^1?T °F?*0CPS °FSTATES °* STATtS''': CH0ICE °F CERTAIN KINDS 0F PROTECTION OR TREATMENT. The following designations are hereby made (please mark the applicable cheek-boxeai: Regional Patent and any other Slate which is a Contracting Sute of the European Patent Convention and of the PCT :—' °* Feneg*U.*TS»:of'"'"' ,MAiM ?"°' Cameroon- Cen,raJ African Republic, Chad. Congo. Gabon. Mali. Maur.unu. and any other Slue which is a Contracting Sute of OAPI and of the PCT; Mother OAPI title desired, specify on dotted line'J >• National Patent (if other kind of protection or treatment desired, specify on dotted line' *>) 1 I [ AT Austria' ' I j KR Republic of Korea''> 3 X' AL' A ustralia* > I | LK Sti Lanka BB Barbados |_j LU Luxembourg'1) 1 ~ BG Bulgaria' ) \~J MC Monaco*1" 1 ~3T BR Brazil' ) i j MG Madagascar Xi CA Canada ^~] MW Malawi'1) . CH and LI Switzerland and Liechtenstein 1 j NL Netherlands ; DE Germany (Federal Republic aty*) .... [X\ NO Norway ! | RO Romania DK Denmark j ] SD Sudan [j ES Spain'1) | SE Sweden j X; Fl Finland |) SU Soviet Union'1) I ; GB Untied Kingdom ~2 HU Hungary X; US United States of America'1). jf JP Japan'1) : j RP Democratic People's Republic of Korea'J' Space reserved for designating States (for the purposes of a national patent) which Have become party to the PCT after the issuance of this sheet. (I) The applicant's choice of the order of designations may he indicated by marking the check-hoses with sequential arable numerals (see also the 'Noies to Box No. V*). (2) The selection of particular States for a Eurooean patent can be made upon entering the national (regional) phase before the European Patent Office (sec also (he "Notes to Box No. V*). (J) - -If another kind of protection or a title of addition or. in the United Stales of America, treatment as a continuation or a continuanon-in-pan Form PCTYRO/IOis desiredI, (seconspecifyd accordin sheet) (Januarg to they instructioniy*0) s given in the "Notes to Boa No. V." See notes on accompanying sheel APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 3 Sheet number. Box No. VI PRIORITY CLAIM (IF ANY). The priority of the following earlier application!si is hereby claimed Country (country-rn which it Filing Date Application Ns Office of Fitinglflll in only :f *ei 'lied if national application: (day. month, year) the earlier application s in one of the countries for which it international ippucatian was filed if regional of interna or a regional application > tional aophcauoa) Hi AUSTRALIA 5 OCTOBER 1990 PK 2658 1*1 Ui • i Letter codes may be used io indicate country and/or Office of fiiing) When the earlier application was filed with the OITiee whieh. for the purposes of the present international application, is the rece-ung Office. ihe applicant may, against paymtnt oftnt nquini ftt. ask the following ?~r the receiving Office is hereby requested to prepare and transmit to the international Bureau a certified copy of the above-mentioned r • earlier application/of the earlier applications identified above by the numbers I insert the applicable numbers i Box No. VII EARLIER SEARCH (IF ANY). Fill m «nere a search iinternational, international-type or atheri by tne International Searching Authority has already been requested lar completed I and the said Authority is now requested to base the international search. to the extent possible, on the results of Ihe said earlier search .Identify such search or request either by reference to the relevant applica tion tor the translation thereof) or by reference io the search request. International application number ai International/regional/national number and country (or regional filing date Office i of other application. Date of request for search Number (if available) given to search request Box No. VIII SIGNATURE OF APPLICANT(S) OR AGENT PHILLIPS ORMONDE & FITZPATRICK Agents for Applicant: rs*-s-^nj> John A. Waters If the present Request form is signed on behalf of anv n~ cam ?y an agent, a separate power of attorney appointing the agent and signed by the applicant is required. If tn such casert i t detired to, -sale i« of a general power of attorney (deposited with the receiving Office i. a copy thereof must be attached to this form. Box No.lX CHECK LIST (To he Client m by tne \-S..CJM. This international application ai filed is accompanied 5y the items checked below- This international application contains the follow ng i.-ser of meets I I • separate signed power of attorney I request -rets ! | | copy of general power of attorney I. description •rets 3 PH pnontydocumentltXteeBoxNo. vii ). claims -ee*.l a i J receipt of the fees paid or revenue stamps i abstract _ •es'.s 5 fx] cheque for the payment of fees S.' drawings -ejtt Teial t4 -rets 6 | j request to charge deposit account • r~] other document (tpeeify) Figaro number ofU»dra«nngs!rian» * digested to accompany the abstract for publication. (The lellewiee » •• M "lied la by Use reeervtstg Office) I. Date of actual receipt of the purported international application 2. Corrected data af actual receipt due to later hut timely received papers or drawings completing the purported international application J. Data of timely receipt of the required corrections under Article 11 of the PCT. 4. Drawings • deceived • No Drawings (Tto Mtotrltii is t* be flUed In by tx* InlirailUael Ba APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 4 ABSTRACT Apparatus for supplying alumina to an electrolysis 5 tank includes a supply chamber in which alumina fed intermittently from a storage container is fluidised by gas injected through a gas-pervious wall- (lo) in the supply chamber beneath the alumina and flows through a discharge conduit (8) to a hole formed in the electrolyte crust. The 10 fluidised alumina flows upwardly to reach the. inlet of the discharge conduit (8) past a hood (7) which prevents non-fluidised alumina from entering the inlet. A method of supplying alumina using the facilities provided by such apparatus is also disclosed. APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 5 METHOD AMD APPARATUS FOR COMTTHTJOnS SUPPLY pp AL-DMIWA This invention relates to a method and apparatus for the continuous supply of alumina or other solid materials to 5 an electrolytic tank in which the alumina is converted to aluminium. In the electrolysis of alumina, solid alumina is dissolved in a tank or pot containing molten electrolyte such as cryolite and it is desirable to maintain the alumina 10 concentration in the electrolyte within a predetermined range. In current practice for the electrolysis of alumina, the alumina is fed in successive doses of predetermined size into one or more holes which are made in the electrolyte crust so that the alumina can be admitted when required. * As 15 the electrolysis of the alumina proceeds continuously, it would be desirable if the alumina consumed in the electrolysis process could be continuously replaced so as to maintain the optimum alumina concentration in the electrolyte. However, the optimum operating conditions are 20 such that the electrolyte crust continuously reforms on the surface of the electrolyte making it difficult to continuously supply alumina to the molten electrolyte beneath the crust. For this reason, known alumina feeding procedures involve the use of a crust breaker which is 2 5 operated intermittently to break the electrolyte crust and form a hole through which the solid alumina can be fed. However, the action of the crust breaker is necessarily such that the crust breaking mechanism, such as a pneumatically operated shaft with an appropriate chisel means (hereinafter APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 6 - 3 - referred to as a plunger) at its free end, will be moved in and out of the hole formed by the plunger. In one known feeding procedure, two separate pneumatic systems are employed, one operating the crust breaking 5 mechanism and the other operating the alumina feeding system. In this procedure, it is possible for the mechanism operating the crust breaking mechanism to form the necessary hole in the electrolyte crust and retract the crust breaker so that the feeding system can then be operated to place a 10 charge of alumina into the hole formed by the crust breaker. In another procedure, a single pneumatic system is used to operate the crust breaking mechanism, and the discharge of alumina from a storage device is co-ordinated with the downward movement of the crust breaker. In this 15 procedure, the alumina charge is thus released when the crust breaker is through the crust so that the alumina is not free to enter the hole in the crust until the crust breaker is retracted. while this procedure has the advantage of a single pneumatic system, it is obvious that 20 not all the alumina will be able to pass through the hole into the electrolyte immediately when the crust breaker is retracted. It is an object of the present invention to provide an apparatus allowing substantially continuous addition of 25 alumina to the electrolyte. The feeder assembly of the present invention is used with a crust breaking mechanism. Any known or other appropriate crust breaking mechanism can be used. The crust breaking mechanism may be pneumatically operated. The crust APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 7 - 4 - breaking mechanism may include a plunger with a cutting edge for breaking the crust mounted on a reciprocable plunger shaft. The feeder assembly further includes at least one 5 storage container comprising a hopper or like vessel for finely divided alumina. Other storage containers may be associated with the feeder assembly for other additives to the electrolysis tank- such as aluminium, fluoride, calcium fluoride, crushed bath, soda ash, or cryolite. The other 10 storage containers may be adapted to feed their contents into the tank in a similar manner to that described below for the alumina. Each storage container is adapted to release a predetermined amount of its contents as required to feed the 15 electrolysis tank. The mechanism employed to measure and release the storage container contents may be any known or other appropriate mechanism. The present invention relates to the supply of alumina from an alumina supply chamber to which predetermined 20 amounts or charges of alumina are fed from the alumina storage container. Accordingly the present invention provides apparatus for supplying alumina to an alumina electrolysis tank including an. alumina storage container adapted to release 25 alumina as required for entry into the electrolyte through a hole formed in crust on the electrolyte by a crust breaking mechanism, characterised in that the alumina storage container is connected to an entry aperture of an alumina supply chamber having a base wall from which inner and outer APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 8 - 5 _ side walls lead towards the upper portion of the chamber to define the entry aperture; a transverse gas-pervious wall extends between the side walls of the chamber spaced from the base wall.and below the chamber inlet to define a plenum 5 into which gas may be introduced through a gas inlet, the gas-pervious wall allowing gas flow sufficient to fluidise alumina but not allowing alumina to enter the plenum; at least one discharge conduit extends through the base wall and the transverse wall to an outlet below the supply 10 chamber, the outlet is positioned in use above the hole formed in the electrolyte crust by the plunger and tihe opposite, inlet, end of the discharge conduit is located above and spaced from the transverse wall; and a hood member is spaced above the discharge conduit inlet but extends 15 below the level of said inlet. The present invention also provides a method for the supply of alumina from a storage chamber adapted for the intermittent supply of predetermined amounts of alumina, characterised in that the predetermined amounts of alumina 20 are fed to a supply chamber, the alumina in the supply chamber is fluidised by gas injection through a gas-pervious wall below the alumina, and the fluidised alumina flows from the supply chamber through a discharge conduit. The alumina supply chamber which forms a 25 characteristic feature of the present invention is partly defined by a base wall and inner and outer side walls which abut or merge with the base wall. The side walls lead from the base wall towards the upper portion of the supply chamber where they may merge towards one another to define APPENDEK 7 PULSE CHUTE PATENT APPLICATION PAGE 9 - 6 - an entry aperture through which alumina fed from the alumina storage container can enter the supply chamber. The supply chamber further includes a transverse gas-pervious wall between the inner and outer side walls. 5 The gas-pervious wall is preferably inclined to the horizontal plane for the reason discussed below. The transverse wall is spaced from the base wall between the base wall and the chamber inlet so as to define, with the base and side walls a plenum into which a gas, 10 preferably air, can be introduced in controlled manner. A gas inlet is provided to enable introduction of gas into the plenum. The gas-pervious material used to form the transverse wall may be any material appropriate to this application. 15 The material concerned must allow a sufficient flow to fluidise the alumina but at the same time should not allow the alumina to enter the plenum. The material should be suitably resistant to the temperatures it will encounter in use. One suitable material is a perforated or woven 20 stainless steel plate. At least one discharge conduit extends through the base wall and the transverse gas-pervious wall of the supply chamber and extends below the supply chamber to an outlet. The discharge conduit outlet is positioned in. use so as to 25 direct alumina into the hole formed in the electrolyte crust by the plunger. The opposite, inlet, end of the discharge conduit is located above and spaced from the transverse wall. The construction of the invention includes a hood member which is spaced above the inlet of the discharge APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 10 - 7 - conduit and extends below the level' of the inlet. The spacing of the hood member and the inlet is such as to prevent gravity flow of alumina from the supply chamber into the conduit as it enters the supply chamber or fills the 5 supply chamber above the level of the inlet. The hood member preferably has a depending skirt, the lower edge of which extends below the level of the conduit inlet. The alumina supply chamber may be a substantially separate vessel located in use to one side of the plunger 10 shaft. In this embodiment, the inner and outer side walls of the supply chamber as referred to above may form a continuous circumferential wall to the chamber. Another form of supply chamber is substantially annular in form and is located concentrically with the plunger shaft. The 15 supply chamber has one or more discharge conduits. In use, the alumina supply chamber of the invention is operated by first supplying a predetermined quantity of alumina from the alumina storage container. The alumina in the supply chamber is then fluidised by injecting gas, 20 preferably air, through the gas inlet to the plenum. The injected gas passes through the gas-pervious transverse wall to fluidise the alumina and allow it to flow past the hood member into the inlet and through the discharge conduit. The gas injection is controlled in relation to its 25 pressure, the time period of injection and the time between successive injections. 3y controlling these parameters for the injected gas in relation to the freguency and amount of predetermined alumina supply to the supply chamber, it is possible to maintain the desired continuity of alumina flow APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 11 - a - from the discharge conduit. Preferably the gas will be injected into the plenum at spaced time intervals to provide fluidising pulses. The pulse length and time between pulses can be varied as required. For example, it is possible to 5 maintain a substantially continuous flow of alumina through the discharge outlet at a rate which is calculated to maintain a supply of alumina within the supply chamber until the release of the next predetermined amount of alumina from the storage container. The flow rate through the discharge 10 outlet from the supply chamber can be varied, depending on the frequency of release of alumina from the stonage container. The flow of alumina through the discharge outlet can be temporarily suspended while the plunger is operated to make or keep clear the necessary hole in the electrolyte 15 crust. The pressure of the gas injected into the plenum may be in the order of five kiloPascal but will vary depending on the permeability of the gas-pervious wall, the height of the alumina above the wall, and the grain size of the 20 alumina which is fed into the supply chamber. PreferabL the gas is injected into the plenum at a pressure and ga flow rate such as to avoid over-f luidisation of the alumina. The gas injection into the plenum can be controlled by any appropriate means, for example 25 solenoid-control on the gas pressure inlet valve can be used to control both the opening and closing of this valve. Alternatively, solenoid-control can be used to open the gas inlet valve against a mechanical pressure applied by a spring or other means tending to keep the valve closed. APPENDED 7 PULSE CHUTE PATENT APPLICATION PAGE 12 - 9 - in order to assist a clearer understanding of the present invention, reference is now made to the accompanying drawing which illustrates one preferred embodiment. It is to be appreciated that this embodiment is given by way of 5 illustration only and that the invention is not to be limited by it. The drawing shows, somewhat diagrammatically, the preferred annular form of alumina supply chamber mounted concentrically with plunger shaft 1 and crust breaking 10 plunger 2. The supply chamber is defined by inner wall 3, outer wall 4 and base wall 5. The upper portion ol outer wall 4 converges inwardly towards inner wall 3 to farm the entry aperture of the supply chamber through which alumina from a storage container (not shown) may enter the supply 15 chamber, diverted, if necessary, by inclined wall 6. Hood member 7 prevents incoming alumina from passing directly into discharge conduit 8. Depending skirt 9 on hood member 7 further, prevents the normal build up of incoming alumina from entering the inlet of discharge outlet 8. 20 Gas-pervious wall 10 defines, with the lower portion of the side walls 3,4 and base wall 5 of the supply chamber, a plenum 11. Gas injected through pipe 12 into the plenum in a controlled manner causes fluidisation of the alumina which has entered the upper part of the supply chamber, thus 25 allowing the fluidised material to flow beneath hood 7, along the body of discharge conduit 8 and through the outlet 13 of this conduit. Outlet 13 is directed towards the hole in the electrolyte crust which has been formed by plunger 2. Gas-pervious wall 10 is- preferably inclined to the APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 13 - 10 - horizontal plane as illustrated. When gas pressure, is applied through the plenum 11, alumina above wall 10 is fluidised but heavier materials which may be present as contaminants in the alumina charge settle to the lower part 5 of the wall 10 and may be cleared periodically. Accumulation of any contaminants at the lower part of the gas-pervious wall minimises the- potential for blockage of the gas flow by these materials and also assists in preventing the contaminants from being carried into the 10 discharge conduit. The present invention makes it possible to extend the time over which a measured charge of alumina is fed into the electrolyte. Thus a relatively large charge can be fed into the electrolyte at an adjustable flow rate and in a 15 substantially continuous manner. The use of relatively large alumina charges reduces the wear on the dispensing equipment and makes more accurate charge measurement possible. APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 14 - ii - 1. Apparatus for supplying alumina to an alumina electrolysis tank including an alumina storage container adapted to release alumina as required for entry into the electrolyte through a hole formed in the crust on the 5 electrolyte by a crust breaking mechanism, characterised in that the alumina storage container is connected to an entry aperture pf an alumina supply chamber having a base wall from which inner and outer side walls lead towards the upper portion of the chamber to define the entry aperture; a 10 transverse gas-pervious wall extends between the side walls of the chamber spaced from the base wall and below the chamber inlet to define a plenum into which gas may be introduced through a gas inlet, the gas-pervious wall allowing gas flow sufficient to fluidise alumina but not 15 allowing alumina to enter the plenum; at least one discharge conduit extends through the base wall and the transverse wall to an outlet below the supply chamber, the outlet is positioned in use above the hole formed in the electrolyte crust by the plunger and the opposite, inlet, end of the 20 discharge conduit is located above and spaced from the transverse wall; and a hood member is spaced above the discharge conduit inlet but extends below the level of said inlet. 25 2. Apparatus as claimed in claim 1 characterised in that the supply chamber is substantially annular in form and is located concentrically with the plunger shaft. APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 15 - 12 - 3. Apparatus as claimed in claim 1 or claim 2 characterised in that the gas-pervious transverse wall is inclined to the horizontal plane. 4. A method for the supply of alumina from a storage chamber adapted for the intermittent supply of predetermined amounts of alumina, characterised in that the predetermined amounts of alumina are fed to a supply chamber, the alumina in the supply chamber is fluidised by gas injection through 10 a gas-pervious wall below the alumina, and the fluidised alumina flows from the supply chamber through a discharge conduit. 5. A method as claimed in claim 4 characterised in that 15 the pressure, the time period of injection, and the time between successive injections of gas are controlled in relation to the frequency and amount of predetermined alumina supply to the supply chamber so as to maintain the desired continuity 'of alumina flow from the discharge 2 0 conduit. 6. A method as claimed in claim 4 or claim 5 characterised in that gas is injected into the plenum at a pressure and gas flow rate such as to avoid 25 over-fluidisation of the alumina. APPENDIX 7 PULSE CHUTE PATENT APPLICATION PAGE 16 APPENDIX 8 A3 U.S.A. PATENT i Appendix 8 A3 U.S.A. Patent Pagel The Commissioner of Patents Q Oh£~ and Trademarks Has received an application for a patent IMted for a new and useful invention. The title and description of the invention are en States closed. The requirements of law have been complied with, and it has been de ~£ termined that a patent on the invention /of) shall be granted under the law >!AnwriC(\^ ne^m »•*• 1 United States Patent Grants to the person or persons having title to this patent theright to exclude WW " - •• others from making, using or selling the Fyy:- •.••:• i invention throughout the United States . ^.;-^, fi of America for the term of seventeen ^>,&'.\; f> years from the date of this patent, sub- i ject to the payment of main tenance fees ip^y as provided by law. Mik l Qimrmssioner of Pattara and Traderntuks u Attest Appendix 8 A3 U.S.A. Patent Page 2 Mill USO05423968A United States Patent m [ll] Patent Number: 5,423,968 Kissane [45] Date of Patent: Jim. 13,1995 [54] ALUMINA SUPPLY APPARATUS FOR 4.431.491 2/1984 Bonny « al 204/67 ELECTROLYTIC SMELTER 4.437,964 3/1984 Gerphagnon et al 204/245 [75] Inventor: James P. Klaaane, Portland. 4463.253 1/1986 Hrinzmarm et al _.. 204/245 X Australia 5.045.168 9/1991 Dalen ei al 204/245 FOREIGN PATENT DOCUMENTS [73] Assignee: Portland Smelter Senrices Pry. Ltd., 92/06230 4/1992 WTPO . Melbourne, Australia [21] Appl. No.: 1M&6 OTHER PUBLICATIONS [22] PCT Filed: Jul. 8,1993 Metallurgy, SU 1560-636-A, p. 23, Apr. 1990. Metallurgy, SU 985-154-A, p. 18 Dec. 1982. [86] PCT No.: PCT/AU93/D0332 Primary Examiner—Donald R. Valentine § 371 Date: Jim. 21,1994 [57] ABSTRACT § 102(e) Date: Jam. 21, 1994 A feeder assembly for an alumina electrolysis cell in [87] PCT Pub. No.: WO94/01fi01 cludes an alumina dose holder (5) defined between inner PCT Pub. Date: Jan. 20,1994 and outer walls (6,7, 7") with an inlet port in the outer wall above an outlet port in the inner wall. The inlet and [30] Foreign Application Priority Data outlet ports (8, 11) are closable and openable by valve Jul. 14. 1992 [AU] Australia PC3496 means formed by relative movement between the outer [51] lot. CL' - C25C 3A4 wall (7, T) and a valve seat (10,12) cooperating with a [52] U.S. CL 204/245; 204/279 seating edge of the outer wall (7, 7"). The valve means [58] Field of Search 204/245, 279 is moved by drive means (13) including a pneumatically operated piston (14) movable within a cylinder (15) [56] References Cited concentric with the shaft (2) of an electrolyte crust- U.S. PATENT DOCUMENTS brealcing plunger (1). The plunger shaft (2) is axially 3.371,026 2/1968 Kitey et ai 204/245 slidable within an annular sleeve (18) of the piston (1*) 3.901.787 8/1975 Niizcki et al 204/245 which is connected to at least one movable component 4,049,529 9/1977 Golla 204/246 of the valve means. 4.321.115 3/1982 Rebmann et al. 204/67 8 Claims, 2 Drawiag Sheets 4,328,085 5/1982 Friedli et al -.... 204/245 Appendix 8 A3 U.S.A. Patent Page 3 U.S. Patent June 13, 1995 Sheet 1 of 2 5,423,968 FIG1 Appendix 8 A3 U.S.A. Patent Page 4 U.S. Patent June 13, 1995 Sheet 2 of 2 5,423,968 FIG2 Appendix 8 A3 U.S.A. Patent Page 5 968 i 2 ALUMINA SUPPLY APPARATUS FOR pneumatic mechanism and the control valve, this neces ELECTROLYTIC SMELTER sary cushioning action affects addition accuracy. In another procedure, two separate pneumatic sys This invention relates to apparatus for the controlled tem5 s ate employed, one operating the crust breaking supply of alumina or other solid materials to an electro mechanism and the other operating the alumina feeding lytic cell in which the alumina is converted to alumi system. In this procedure, it b possible for the mecha num nism operating the crust breakmg mechanism to form In the electrolysis of alumina, solid alumina is dis the necessary hole in the electrolyte crust and retract solved in a tank or pot containing molten electrolyte 1° the crust breaker so that the feeding system can then be such as cryolite and it is desirable to maintain the alu operated to place a charge of alumina into the hole mina concentration hi the eJectrolyte within a predeter formed by the crust breaker. There is less air usage as mined range. In current practice for the electrolysis of the dosing cylinder of the alumina feeding system is of alumina, the alumina is fed in successive doses of prede smaller capacity and may operate more often than the termined size into one or more holes which are made in crust breaker as it is independent of the crust breaker the electrolyte crust so that the alumina can be admitted mechanism. However, this known procedure requires when required. As the electrolysis of the alumina pro separate housings for the crust breaking mechanism and ceeds continuously, it would be desirable if the alumina the alumina feeding system. These separate housings consumed in the electrolysis process could be continu not only reduce the space available for operating above ously replaced so as to maintain the optimum alumina the electrolysis tank but also complicate the construe - concentration in the electrolyte. However, the optimum i tion of the whole assembly. Hence this design is more operating conditions are such that the electrolyte crust expensive than the single pneumatic mechanism design. continuously reforms on the surface of the electrolyte, In our prior Australian patent application no. PK making it difficult to supply alumina continuously to the 2658/90 (which forms the basis for International Appli molten electrolyte beneath the crust. cation PCT/AU91/00169) we proposed a feeder assem- For this reason, known alumina feeding procedures ; bly in which a valve mechanism concentric with the involve the use of a crust breaker which is operated shift of the crust breaking plunger is operable in re mtertnittently to break the electrolyte crust and form a sponse to the initial part only of the downward move hole through which the solid alumina can be fed. How ment of the plunger but achieves the advantage of hav ever, the action of the crust breaker is necessarily such ing the plunger out of the alumina flow. However, like that the crust breaking mechanism, such as a pneumati t the first design, it is affected by cushioning, as the crust cally operated shaft with an appropriate chisel means breaker cylinder is much larger than necessary for dos (hereinafter referred to as a plunger) at its free end, will ing of the alumina. This design is still affected by high be moved in and out of the hole formed by the plunger. air usage but its accuracy of dosing is less affected by In one known feeding procedure, a single pneumatic speed. mechanism is used to operate the crust breaking mecha i It is therefore an object of the present invention to nism, and the discharge of alumina from a storage de provide an improved alumina feeder assembly which vice is co-ordinated with the downward movement of enables direct feed of alumina into the hole (as is the the crust breaker. In this procedure, the alumina charge case with the two separate pneumatic systems design is thus released when the crust breaker is through the and our Australian patent application PK 2658/90) and crust so that the alumina is notfree to enter the hole in > significant reduction in air usage compared to a design the crust until the crust breaker is retracted. While this using a single pneumatic drive cylinder. procedure has the advantage of a single pneumatic sys Accordingly, the present invention provides a feeder tem, it is obvious that not all the alumina will be able to assembly for an alumina electrolysis tank including a pass through the hole into the electrolyte immediately crust breaking mechanism operable to break a hole in when the crust breaker is retracted. > crust formed on the surface of molten electrolyte, the It will also be apparent that with this form of mecha crust breaking mechanism including a plunger with a nism, the plunger must travel through the crust each cutting edge mounted on a reciprocable plunger shaft time a charge of alumina is to be introduced into the and an alumina storage container adapted to release electrolyte. This not only involves use of sufficient air alumina as required into a dose holder, characterised m to drive the plunger but also may involve dipping the ) that the dose holder is defined between inner and outer plunger into the electrolyte with each stroke. It is desir walls, an inlet port is formed in the outer wall above an able to reduce the number of times the plunger contacts outlet in the inner wall whereby alumina can flow the electrolyte as far as possible so that wear of the through the dose holder from inlet port to outlet port plunger can also be reduced. under the influence of gravity, the inlet and outlet ports Accurate alumina flow control is made difficult by 5 being cVosable and openabfe by valve means formed by the required relationship between the plunger move relative movement between the outer wall of the dose ment and movement of the flow control valve control holder and a valve seat which cooperates with a sealing ling alumina discharge. It will be understood that the edge of the outer wall, the valve means being movable force available and speed necessary for plunger move by drive means including a pneumatically operated ment must be sufficient to achieve crust breakage. This 3 piston movable within a cylinder concentric with the force, speed and the amount of air needed to produce plunger shaft, the piston having an annular sleeve axi- them are far greater than that needed to move the alu ally slidable within the cylinder and the plunger shaft mina flow control valve and it is therefore necessary to being axially slidable within the sleeve which is con cushion the final stages of the pneumatic mechanism nected to an extension sleeve in turn connected to at against the considerable force available for plunger 5 least one movable component or the valve means. movement. Because of the interaction between the The feeder assembly of the present invention includes a crust breaking mechanism which is preferably pneu matically operated. The crust breaking mechanism m- Appendix 8 A3 U.S.A. Patent Page 6 968 3 5'423' eludes a plunger with a cutting edge for breaking the 4 crust, mounted on a reciprocable plunger shaft. savings in air usage compared with a single pneumatic The feeder assembly may be associated in use with at mechanism. least one storage container comprising a hopper or The apparatus of the invention preferably includes a similar vessel forfinally divide d alumina- Other storage 5 housing which surrounds the plunger shaft and the containers for other additives to the electrolysis tank. generally concentric feed mechanism components de such as aluminium fluoride, calcium fluoride, crushed scribed above. The lower end portion of the housing bath, soda ash. or cryolite may be associated with simi preferably tapers inwardly to facilitate direction of the lar feeder assemblies. The other storage containers may alumina falling towards the electrolysis tank after leav thus be adapted to feed their contents into the tank in a 10 ing the outlet port of the feeder assembly. The lower similar manner to that described below for the alumina. edge of the housing thus defines the periphery of a The storage container or containers are adapted to release port through which altnnina leaves the housing feed their contents as required into a dose holder. Pref as it fails towards the tank erably the storage container contents are first fed into a Alternatively, the lower end portion of the housing supply chamber which has an exit port communicating 15 may be formed to provide two or more downwardly with an inlet port of the dose holder. The dose holder is tapering outlets directed towards the hole formed in the defined between inner and outer walls which are prefer electrolyte crhst by the plunger. ably concentric with each other and with the plunger To assist a further utivderstandtrig of the invention. shaft. The inlet port of the dose holder is formed in its reference is now made to the accompanying drawings outer wall. The dose holder farther includes an outlet 20 which illustrate one preferred embodiment of the pres port located below the inlet port so that alumina or ent Invention, It is to be appreciated that this embodi other material can flow through the dose bolder from ment is given by way of illustration only and that the inlet port to outlet port under the influence of gravity. invention is not to be limited by h. Both the inlet port and the outlet port are closable BRIEF DESCRIPTION OF DRAWINGS and openable by valve means formed by relative move- 25 The drawings show, somewhat diagrammatically. ment between the outer wall of the dose holder and a sectional views of a preferred form of feeder assembly. valve seat which co-operates with a seating edge of the In FIG. 1, the dose holder is open to the entry of outer wall. Preferably the inlet port is closed by a mov alumina. able upper wall seatable in a fixed upper seat and the FIG. 2 shows the dose holder at the opposite extreme outlet port is closed by movement cf a movable lower 30 of the valve movement, closed to the entry of alumina seat into abutment with a fixed lower wall. but open to discharge alumina. In accordance with the invention, we provide for DESCRIPTION OF PREFERRED EMBODIMENT movement of the valve means by a drive means includ In the drawings, a crust breaking plunger 1 is carried ing a pneumatically operated piston movable within a by plunger shaft 2. A storage container (not shown) cylinder concentric with the plunger shaft. This piston 35 feeds alumina or other tank additive as indicated by includes an annular sleeve axiaily slidable within the arrows A into an annular supply chamber 3. Supply cylinder and the plunger shaft is axiaily slidable within chamber 3 has an exh port or ports 4 which remain open the sleeve. The piston sleeve is connected to a prefera allowing the alumina to fall towards annular dose bly annular extension sleeve which is in turn connected holder 5. Dose holder 5 is defined between an inner wall to at least one movable component of the valve means. 40 6 and an outer wall 7, T, both of which are concentric In one preferred construction, the drive means is with plunger shaft 2. connected to move the upper wail of the dose holder The outer wail of the dose holder is formed in two and the movable lower seal simultaneously. Thus the parts. 7 and T. In the illustrated embodiment, the upper outlet port is closed as the inlet port is opened. Con part 7 of the outer wall is movable and in its raised versely, the inlet port is closed as the outlet port is 45 position provides an inlet port 8 to the dose holder S opened. between its lower seating edge 9 and valve seat 10. The Alternatively, the construction may provide for inde movement of outer wall part 7 and its association with pendent operation of the inlet port valve means and of valve seat 10 thus provide a valve means for inlet port the outlet port valve means. For example, the piston 8. may be connected to move the inner wall and the valve 50 A valve means for outlet port 11 of the dose holder is seat associated with the outlet port of the dose holder, provided by the relative movement between movable and the outer wall of the dose holder is connected to the valve seat 12 and thefixed lowe r wail part T of the dose cylinder wall which is movable concentrically with the holder outer wall. piston. The movable cylinder and outer wall control the Movement of the respective valve means is con operation of the valve means associated with the inlet 55 trolled by drive means 13 which includes a pneumati port. cally operated piston 14 movable within cylinder 15. Whatever particular construction is used, the use of Air supply lines 16 and 17 are used to activate the up the plunger shaft concentrically slidable within the ward and downward movement of piston 14 within piston provides valve means operable independently of cylinder 15. The lower end of piston 14 is connected to the plunger. Accordingly, the valve means can be oper- 60 extension sleeve 18 which in turn carries the upper part ated as often as required to add alumma to the electroly 7 of the dose holder outer wall and the lower wall 15 of sis tank while the plunger needs to be operated only the dose holder which includes movable valve seat 12. when necessary to break the crust and allow access of As will be appreciated, downward movement of pis the alumina to the electrolyte mix. The cylinder driving ton 14 from the position shown in FIG. 1 brings the the valve means needs only to be of considerably 65 sealing edge 9 of part wall 7 into contact with valve seat smaller stroke than that driving the crust breaker. The 10, thus closing dose holder inlet port 8. At the same cylinder driving the valve means can also be of smaller time the downward movement of piston 14 separates piston area. There are thus made possible significant Appendix 8 A3 U.S.A. Patent Page 7 5,423,968 5 6 movable valve seat 12 from the lower edge 21 of dose a sealing edge of the outer wall, the valve means being holder part wall T, opening outlet port 11 of dose movable by drive means including a pneumatically holder S and allowing the alumina within the dose operated piston movable within a cylinder concentric holder to flow out under the influence of gravity. This with the plunger shaft, the piston having an annular downwardly flowing alumma ts directed by the in- 5 sleeve axiaily slidable within the cylinder and the wardly tapermg lower part of hotaing 2C'towards the phlnger shaft ^ MiMy sljdable withb me d plunger artd plunger^ft^tnd thus towards the hole ,n which « connoted to an extension sleeve in turn con- the crust which has been broken by the plunger ^^ to at least OM moViWe CQ As drive means 13 for the valve means of dose holder means. 5 is operable mder^nderitly of the drive means (not 10 x A fceder ma^ „ ^^ m ^ I^*S. ^g,.,\v iT^l^i ^ W ^ m **the inner "** outer wal,s of the dose holder finer control than that achievable if the plunger move- afe ^^^ ^ ^ other ^^ ment and dose holder valve means movement are inter- sh^n dependent and is of smaller area and stroke to the crust •» Afi.«^», ™,.„*~.i» i , „i • ~a • « . i_ . , . s. w 3. A feeder assembly as claimed m claim 1 character- breaker and hence saves air usage. 15 - . - ... ,1. • . . _ • , ,. >-i«a«»tw T» - c _.u J flT^ i_ lset* w1that 'he inlet port is closed by a movable unper It is a further advantage of the present invention that „,„„ „„,_ui„ - = \, . * . , uFt«;i ., , ., s„^5._j s J _. -s - . - . wa« seatable in afixed uppe r seat and the outlet port is the mecMnism can be fitted relatively easily mta exist- . ., _ . » 1.1 , . y . • s_- u 1 1: J - IJ closed by movement of a movable lower seat into abut- i»g plants which use a single pneumatic drive cylinder, «.„-„ ft. as J 1 n «»*-«. amu owi, e i , _ . r .. • - , . - . , - ment with a fixed lower wall. when replacement of the original drive and alumma . . ,„ .„ _., , . , . „ , delivery mechanism is considered inxessary or desir- 20 * A ^."T*** " dauned m da™ 3 characte_r- able. This can lead to savings in space and in costs asso- "** *•» *«'**? "^ B "ff ^ W ^ ** ciated with the structure requiredto suppor t the feeder """I wf * «**,dose hoWer ««• ** movable lowM assembly. Costs can also be reduced by the use of this « sJn"lltan«ous,y- invention to reduce the number of plunger movements . * A <«der assembly as claimed in clam 3 character- and plunger wear. The addition of alttmina indepen- „««.»•-« the drive means ts connected for indepen dently of plunger movement also allows the possibility dw" ™vaaeat <*** respective inlet port valve means of more frequent ahrmina additions, approaching con- ""* ""j" P°n va,vc, racanf . dnuous addition of alumina. The abilityto operate the . 6- A feeder "»«--bt*, as claimed m claim 5 character- respective valve means of the inlet and outlet of the wed m that the piston is connected to move the inner dose holder independently further allows greater accu- 30 wal1 *nd lhe vaIve f31 "weiiied with the outlet port racy of the alumina additions by preventing alumina and the °uter "^ ofth e ^^ ^lder » connected to the from entering the dose holder as the dose holder con- cylinder wall which e> movable concentrically with the tents are released. piston to control the operation of the valve means asso- I claim: ciated with the inlet port. 1. A feeder assembly for an alumina electrolysis cell 35 7-A feeder assembly as claim in claim 1 chiiracterised including a crust breaking mechanism operable to break m that the dose holder is concentric with the plunger a hole in crust formed on the surface of molten electro- **"&• and a housing surrounds the plunger shaft and lyte, the crust breaking mechanism including a plunger dose holder, the lower end portion of the bowing taper- with a cutting edge mounted on a reciprocable plunger ing inwardly with its lower edge defining the periphery shaft, and an alumma storage container adapted to re- 40 of a release port though which alumina leaves the hous- lease alumina as- required into a dose holder, character- rng and falls towards the tank. ised in that ihe dose holder is defined between inner and 8. A feeder assembly as claimed in claim 1 character- outer walls, an inlet port is formed in the outer wall ised in that the dose holder is concentric with the above an outlet in the inner wall whereby alumina can plunger shaft and a housing surrounds the plunger shaft flow through the dose holder from inlet port to outlet 45 and dose holder, the lower end portion of the housing port under the influence of gravity, the inlet and outlet 55 being formed to provide two or more downwardly ports being closable and openable by valve means tapering outlets directed towards the hole formed in the formed by relative movement between the outer wall of electrolyte crust by the plunger. the dose holder and a valve seat which cooperates with * * * * * 50 60 Appendix 9 Diary of Changes Page 1 APPENDIX 9 DIARY OF CHANGES Line 2 Start Up: 15.10.86 - 01.07.87 all bar 6 pots 27.02.88 Line 1 Start Up: 07.03.88 - 23.09.88 Fail Safe dates: Start - Terry 01.03.87 -Parker 01.07.88 Retrofit -Parker01.06.89 Dec 88 All Terry piston rods being replaced by Kempe rods. 04.05.90 Commenced measuring plunger OD at overhaul. 05.05.90 Commenced tieing up feeders after overhaul for ease of installation. 06.05.90 Initial design A2 feeders. 28.05.90 Commenced using code on plungers (25E90) and tagging failed feeders. 14.06.90 Huck visit for bolting feeders - unsuccessful. 15.06.90 Commenced discussions for A2 patent. 25.06.90 Commenced CETEC R & D on plungers - examine failed plungers and investigate ceramics. 28.06.90 Visit Parker and commissioned Unisearch to assist in R & D and advise on cushioning analysis of cylinders. 27.07.90 Commenced R & D on spring design at Kempe (Geelong). 13.08.90 Commenced using reinforced spools and not using old pipe in spools. 22.08.90 Sourced marinate and Unasco 808 for insulation band and SG-200 insulator for bolts. Oven tests commenced. 31.08.90 Commenced 35mm steel washers for feeder bolts. 13.09.90 High temperature die steel springs tested OK. Ordered 500. 15.09.90 Commenced using cast plungers with pin area increased (to stop pin being caught on super chute) and straight sides (to stop dragging on lip halfway up). 25.09.90 Commenced using Bakelaque washers on bolts of supers. 04.10.90 Applied for provisional patent for A2 patent. 22.10.90 Commenced changing piston seals only on failure. 05.11.90 Started using straight plungers. 19.11.90 Tested A2 in off-line pot - chute hit by anode. Redesign. 20-21.11.90 Qishioiiing test at Parker - high deceleration rate 50G (500m/sec) versus 25G target. 27.11.90 Cushion tests at Parker - confirmatory test show same results on old and new cylinders. 15.12.90 Terry found incorrect piston seal width - changed from 25 to 50 thou. 15.12.90 Started usingflat bottome d cast plungers to make bigger hole. 01.02.91 Started using trial 200 Unasco 808 insulation bands. 04.02.91 A2 feeders (double chute dribble feed) installed in 4070 and single chute in 3092 (AEDD feeder). 05.02.91 Repair failure correct on history for all parts after this date. 07.02.91 Traced "not feeding" to seized spool 146mmOD better than 152mmOD. 12.02.91 Installedfirst pot (1011) with trial plungers. 15.02.91 Started changing out 152mmOD spools to 146mmOD spools. 22.02.91 Silicon carbide plungers removed due to breakage. 05.03.91 3092 and 4070 changed to 50° double chutes (Dribble feed). 07 03 91 4033 changed to 5 inconel 65mmOD plungers. 22.03.91 12 and 15mm nozzle placed in chutes of 4070 and 3092 (50% feed rate). 24.03.91 3024 changed to 5 long cast plungers. 25.03.91 Started tests on cylinder cushioning at Kempe. 27.03.91 Trialled belville washers on Parker cylinder to improve cushioning. Not successtui. 03.04.91 2061 #5 chute changed to dual outlets (nozzles installed 10.04.91). Appendix 9 Diary of Changes Page 2 17.04.91 Meeting IXL on new specification of cast iron plungers. 26.04.91 60 cast iron plungers made as Stage 1 of plunger trials - gauge capabilities and location of hardness point for measurement. 29.04.91 International patent application for A2 feeder and airslide chute. 01.05.91 Visit ITC Wollongong re angles of chutes for reacted ore. 03.05.91 Nozzles taken off 3092 - blocking. 09.05.91 Started testing Terry cylinders on cushion rig.at workshop 14.05.91 Feeder survey complete - all locations have cylinder number identified - 18.05.91 Rest of nozzles out of 4070. 23.05.91 Started tests on slowing feeders on 8 pots. 01.06.91 Started testing breaking hole at zero speed 03.06.91 Started fracture tests on piston/shaft connection at Unisearch. 12.06.91 Tested new Parker cylinder insulation 30M OHM to 1.5G OHM. No insulation on old cylinders 20.06.91 Tested speed of cylinders with video. 22.07.91 Completed piston rod/shaft fracture tests - Terry better than Parker. 25.07.91 Training started on FRS for tradesmen at Kempe. 25.07.91 Started using high temperature springs. 02.08.91 1001-1051 dwell time 2.5 to 2.0 sees. Sep 91 Stopped spigotted flanges.as not making a difference 01.09.91 Report "Pot Alumina Feeder Design Review" published. 28.09.91 Visit Brenko Thermospray for plunger coatings 02.09.91 Started using studs on Parker broken piston rods. 17.09.91 All pots 2.0 sec dwell time. 17.09.91 Changed Terry rod and piston seals to blue teflon in white cartridge. 18.09.91 2065, 2066, 2067 and 2068 dual dwell time trials. 07.10.91 3026 superstructure modified with holes for better feeder venting. 15.10.91 Reduced grease on seals on overhaul. 16.10.91 20 Terry prototype teflon rod seal in service. 23 10.91 Test insulation on feeder in workshop - 0.4G new and zero old. Insulation breaks down in service. 28.10.91 At Unisearch (noise reduction for silencers, fatigue tests, orifice design). Nov,91 Parker rod seals changed from 530 (Parker) to 548 (Terry). Terry changed from 432 to 442. 12.11.91 Line 1 now slow speed. 12.11.91 Started Terry FEC in Parker. 13.11.91 Started 103E slow speed 3.0 sec dwell time. 19.11.91 Future Parker piston rods 4140 5 thou (60 micron) chrome from 10L45, 2 thou (25 micron). 19.11.91 Purchased cast double outlet chutes. 22.11.91 Purchased 12 new feeder racks with safer access. 25.11.91 All pots (both lines) at slow speed. 02.12.91 Line 1 4 sec to 3 sec dwell time and Line 2 @ 4 sec dwell time. 05.12.91 New pot 1075 is A2. 28.01.92 First Portland Aluminium spec, plungers in service - still using HR for 1 month. Feb 92 Changed to blue FEC casings and blue bushes for Terry. 01.02.92 Started using 4140 Parker piston rods. 02.02.92 Started using bronze inserts instead of hardened inserts. 26.02.92 R&D Parker teflon rod seals arrived. 02.03.92 3049 cast changed to A2 feeders and IXL cast chutes. 27.03.92 Cylinder trial start at Parker. 01.04.92 Report "Optimising Pot Feeders" published. 08.04.92 At Rockdale and Wenatchee to give presentation on feeders to Feed Alumina NetworK. 27.05.92 Insert on all feeders (81mm stroke) ex workshop. 08 06.92 1064 A2 changed to AEDD - dags causing anode effects. 12.06.92 3095 1,3,5 A2 feeders changed to AEDD - dags causing anode effects. 24.06.92 At Parker for autopsies on cylinder trial. Appendix 9 Diary of Changes Page 3 01.07.92 Parker piston back up washers no longer used. 28.07.92 Trial single outlet chute 2009 # 1. 07.08.92 Changed dual dwell time short circle from 1.7 to 1.9 sec (100 pots). 17.08.92 Started routine shotsize testing of new pot feeders. 02.09.92 Started 53mm stroke spool inserts on all feeders ex workshop. 03.09.92 Started stroke testing of feeders in pots - pressures and stroke times. 30.09.92 A2's changed to 2 sec short and 4.5 sec long every 3rd cycle. Sep 92 Started flange insulation and boomerangs. 06.10.92 20 Norton seals arrived. 21.10.92 2009 chutes changed to single outlet. 25.10.92 20 Dover seals arrived. Nov 92 Changed A2 pots to single outlet chutes. 21.12.92 All pots 4.5/2.2/2 dwell time. 24.12.92 Changed or cut off #3 feeder plus others on all A2 pots as dag getting caught and chutes - excess anode effects. 29.12.92 5.0/2.0/2 dwell time. 30.12.92 3095 changed from A2 to AEDD - dags. 06.01.93 Started installing Parker piston back up washers again. 06.01.93 High feeder changeouts due to pointed feeders - 20 feeders had only plungers changed (no cylinder overhaul). 15.01.93 Started quick overhauls based on air leaks. 18.01.93 All pots dwell 4.5/2.0/2 -> 4.5/2.2/2. 02.02.93 100 trial rubber shirts on assemblies to prevent ore leaks. 16.02.93 Accepting air leaks of under 30 1/min 16.02.93 Don't overhaul cylinders if plunger worn only. 18.02.93 Dual dwell time back on all pots. 22.02.93 Trial mobile plunger coolers operating. 22.02.93 Started trial of seals/plungers/bushes on 9 new pots. 26.02.93 Increased grease. 26.02.93 First Bimetal plungers installed. 01.03.93 Report "Point Feeding Literature Survey" published. 05.03.93 4011 changed to double outlet chutes - not dribble feed. 05.03.93 Purchased 40 air leak testers.to measure air leaks in pots. 14.03.93 Started work on plunger penetration investigation. 19.03.93 Started 8 day trial of plungers in liquid bath. 19.03.93 Started using studs for broken piston rods again - 360 days extra life. 27.03.93 Started first 7 pots using plunger coolers to stop dagging on start up. 29.03.93 Traced 24 broken plungers - 70mm shorter and 100% better wear rate than feeders in same pot. 29.03.93 Commenced national phase PCT patent applications for A2 feeders. 01.04.93 No more overhauled feeders installed on new pots. 06.04.93 10 feeders installed with 50mm shorter plunger. 08.04.93 Purchased 50 AEDD and 35 A2 feeder chutes. 10.04.93 Started cone lock nuts on boomerang brackets. 20.04.93 Visit from Huck re better feeder bolting - rej ected. 22.04.93 First superstructure modified with feeders not removed by operators (1066). 23.04.93 Purchased 35 non magnetic slings. 26.04.93 Alternating plungers (PA, HR, Incoloy, 310SS, 21%CR, 25% CR, 27%CR, Ni-resist) for all overhauls until used. 28.04.93 Installed short plungers on all feeders in 1031 and 4023. 30.04.93 Engaged CETEC, Shell and Optimol for grease investigation. 05.05.93 Started al. bronze for hardened inserts. 07.05.93 All pots 5 sec dwell time after set for 2 hours. 10.05.93 Started stepped washer welded to bolt on top flange. 10.05.93 Started installing swivelfittings in centre port. 12.05.93 Started changeout of A2 chutes to double oulet (2069). 26.05.93 Started installing short plungers in another 7 pots. Appendix 9 Diary of Changes Page 4 02.06.93 Started installing double outlet chutes on 10 AEDD pots 03.06.93 7 pots with short plungers completed. 10.06.93 1096 installed with split air header. 28.06.93 All west half of pot lines covering shift after set - 80% reduction in blocked feeders. July 92 Air leak and plunger survey completed. 01.07.93 Commenced feeder removal for inserts and west end ownership change. High turnover of feeders started. 01.07.93 Commissioned and trained operators on FRS. 01.07.93 All plungers now 50mm shorter ex Kempe. 01.07.93 All magnalube finished. Now using TRANZ 414. 07.07.93 Patent applicationsfiled fo r A3 feeders (PCT plus 7 others). 12.07.93 All A2 now converted to double outlet chutes. 13.07.93 A2 feeders changed 4.0/2.2/3 to 4.5/2.2/3 dwell time. 15.07.93 Purchased 2600 new feeder ID tags and 2 extra Terry hollow cylinders for A3. 15.07.93 3002-16 even numbers changed to break every 3rd cycle (not 2nd cycle). 15.07.93 First batch of 310SS cast production model ex DCL. 16.07.93 Stopped reusing feeders in new pots. 19.07.93 All slings changed to non magnetic. 21.07.93 2064 converted to A2 feeders. 21.07.93 Finished installing double outlet chutes on 10 AEDD pots. 22.07.93 Plunger shaft acceptance changed from 47mm to 45mm. 26.07.93 Increased grease quantity to about lOOg/overhaul. 26.07.93 Introduced "alumina in assembly" fault. 28.07.93 Engaged Unisearch to fatigue test A3. August,93 Trimming guide to prevent partial vacuum in spring chamber. 03.08.93 Started cutting guides for less suction in spring chamber. 04.08.93 Restrictors out of 2018-34 odd numbered pots. 09.08.93 103 odd 4/2.7/2; even 5/2.2/2 dwell time trials. 13.08.93 2073#3 first A3 installed (with kidney plate turned off) -Ortman. 13.08.93 All slings new design. 13.08.93 102 all 4.5/2.4/2 18.08.93 First A3 removed due to incorrect air set up - 8 second dwell and large dags. Crust break rod seal failed and minor air leak on Ortman feed cylinder. 23.08.93 103 all 4.5/2.4/2 dwell time. 27.08.93 101, 103, 104 all on 4.5/2.4/2. Sept,93 TA2R2, GPSO, A2M grease trials started (alternating feeders). Sept,93 Started drilling holes in assembly to prevent partial vacuum in spring chamber. 08.09.93 Started drilling holes in assemblies to depressurise spring chamber and to assist cleaning. 14.09.93 40 stainless spanners for feeder changeouts arrived. 20.09.93 Ordered Kennett ladder for feeder changeouts. 18.10.93 2073 #3 A3 feeder removed for inspection and replaced with Terry feed cylinder. Crust break seals, bad leak but minor on feed cylinder. Bushes show only pitting from gas attack - 60,000 cycles equivalent to 80 days operation. 28.10.93 20 Atlas Copco cylinders arrived. 28.10.93 4011 (A2 pot) taken off line. Nov,93 Parker rod seals changed from 548 to 548s (short reversed with teflon wiper). Terry changed from 442 to 430Y to same design. 08.11.93 Changed to wider walled FEC. 10.11.93 1094 on-line as split header. 12.11.93 1022 on split header. 01.12.93 2070 on A3 indep feed. 07.12.93 2070 on end stroke sensing. 10.12.93 Started installing Nicrofer plungers. 30.12.93 # 1 and 3 removed from 2070 - leak and siezed. 01.01.94 Normal bolts on boomerangs. 25.01.94 Trial Kennet ladder arrived. Works but heavy to move. Appendix 9 Diary of Changes Page 5 01.02.94 2046 #1 and 2047 #3 changed to 155mm long plungers (50mm short). No problems after 6 weeks operation before pot cut out. 08.02.94 Inspection of Ortman removed ex 2070 #3 on 31.12.93 - lock lite 567 siezed piston rod. 08.02.94 2070 #1 Terry out - ore leak due to non-viton 'O' ring on FEC. 08.02.94 2073 end stroke sensing installed. 11.02.94 2070 #1 seized (Ortman) no fault found at room temperature. Non-conforming Kempe put in. 11.02.94 2070 charged to feed every time. 15.02.94 2070 #1 too slow, orifice removed. 21.02.94 2070 #2 not feeding - anode effects. 28.02.94 2074 end stroke sensing commissioned. 01.03.94 Shot size testing on 2070 showed 2.5s dwell time caused variable shot size. Changed to 4.0s. Anode effects dropped from 2/day to 0.1/day. 07.03.94 2075 sequential feed installed but not commissioned. 18.03.94 Reinstalled all restrictors. All pots running at same speed. Result of trials from 8/93 to 3/94 of alternating sections of restrictors in and out. Statistical evidence of less anode effects at slower speed. June,94 Using TA2R7 grease. APPENDIX 10 CONFIDENTIAL TO ALCOA *** PLANT CODES INTEGRATED SODERBERG T1 Portland (Alcoa) S1 Lista (Elkem) T2 British Alcan (Alcan) T3 Mosjoen (Elkem/Alcoa) T4 Le Terriere (Alcan) T5 Grande Baie (Alcan) T6 Pt Henry (Alcoa) T7 Sao Luis (Alcoa) GUILLOTINE T8 Boyne (Comalco) T9 Wenatchee (Alcoa) G1 Bell Bay (Comalco) T10 Kurri Kurri (Alcan) D11 N2AS (Comalco) T11 Massena (Alcoa) T12 Mead (Kaiser) T13 Badin (Alcoa) T14 Seebree (Alcan) T15 Rockdale (Alcoa) T16 Warrick (Alcoa) T17 Mt Holly (Alumax) INDEPENDENT D1 ISAL (Alusuisse) D2 Karmoy (Hydro Aluminium/Pechiney) D3 Granges (Reynolds) D4 Alouette (Pechiney) D5 Dunkirk (Pechiney) D6 Beconcour (Pechiney) D7 Aluar D8 Lauralco (Pechiney) D9 Baie Comeau (Pechiney) D10 Tomago (Pechiney) D11 Ti Wai (NZAS) D12 Tennessee (Alcoa) D13 Ardal (Hydro Aluminium) D14 St John de Maurienne (Pechiney) D15 Lockerby (Pechiney) D16 Alba (Pechiney) D17 VAW D18 Venalum (Hydro Aluminium) D19 Kidricevo (Pechiney) D20 Sundalsora (Hydro Aluminium) D21 Zia (Pechiney) D22 LMG (Alusuisse) D23 Steg (Alusuisse) D24 Inalco (Pechiney) D25 Alusaf (Pechiney) APPENDIX 10 *** CONFIDENTIAL TO ALCOA *** PLANT CODES INTEGRATED SODERBERG T1 Portland (Alcoa) S1 Lista (Eikem) T2 British Alcan (Alcan) T3 Mosjoen (Elkem/Alcoa) T4 Le Terriere (Alcan) T5 Grande Baie (Alcan) T6 Pt Henry (Alcoa) T7 Sao Luis (Alcoa) BARBREAK (GUILLOTINE. T8 Boyne (Comalco) T9 Wenatchee (Alcoa) G1 Bell Bay (Comalco) T10 Kurri Kurri (Alcan) D11 NZAS (Comalco) T11 Massena (Alcoa) T12 Mead (Kaiser) T13 Badin (Alcoa) T14 Seebree (Alcan) T15 Rockdale (Alcoa) T16 Warrick (Alcoa) T17 Mt Holly (Alumax) INDEPENDENT D1 ISAL (Alusuisse) D2 Karmoy (Hydro Aluminium/Pechiney) D3 Granges (Reynolds) D4 Alouette (Pechiney) D5 Dunkirk (Pechiney) D6 Beconcour (Pechiney) D7 Aluar D8 Lauralco (Pechiney) D9 Baie Comeau (Pechiney) D10 Tomago (Pechiney) D11 TiWai(NZAS) D12 Tennessee (Alcoa) D13 Ardal (Hydro Aluminium) D14 St John de Maurienne (Pechiney) D15 Lockerby (Pechiney) D16 Alba (Pechiney) D17 VAW D18 Venalum (Hydro Aluminium) D19 Kidricevo (Pechiney) D20 Sundalsora (Hydro Aluminium) D21 Zia (Pechiney) D22 LMG (Alusuisse) D23 Steg (Alusuisse) D24 Inalco (Pechiney) D25 Alusaf (Pechiney) 1^- Ctf)N t c T3 CM 8 > > > > > > > > r> <£ Si- 4fc CO —- .—. ^ JZ CN tf)^» ^ CN J" o o o o o o o o o UJ Q S-- *; * S CO .2rn ^_IS P TJ to 2 2 2 2 2 > ro - to P o O) H z z z z fc a '> ^CO CD<» n £ O) < CM CO ID d> 6 *— CO 4 CO < co jz! 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