U.S. Department of Transportation Publication No. FHWA-NHI-16-009

Federal Highway Administration FHWA GEC 012 – Volume II 

July 2016

NHI Courses No. 132021 and 132022

Design and Construction of

Driven Pile Foundations – Volume II 

Developed following:

AASHTO LRFD Bridge Design Specifications, 7th Edition, 2014,with 2015 Interim.

AASHTO LRFD Bridge Construction Specifications, 3rd Edition, 2010, with ‘11, ‘12, ’13, ‘14, and ‘15 

Interims.
NOTICE

The contents of this report reflect the views of the authors, who are responsible for the facts 

and accuracy of the data presented herein. The contents do not necessarily reflect policy of 

the Department of Transportation. This report does not constitute a standard, specification, or 

regulation. The United States Government does not endorse products or manufacturers. Trade 

or manufacturers′ names appear herein only to illustrate methods and procedures, and are 

considered essential to the objective of this document. 
Technical Report Documentation Page

1. REPORT NO.

FHWA-NHI-16-010

2. GOVERNMENT ACCESSION NO. 3. RECIPIENT'S CATALOG NO.

4. TITLE AND SUBTITLE

Geotechnical Engineering Circular No. 12 – Volume II 

Design and Construction of Driven Pile Foundations

5. REPORT DATE

July 2016

6. PERFORMING ORGANIZATION CODE

7. AUTHOR(S)

Patrick J. Hannigan, PE, Frank Rausche, PhD, PE, 

Garland E. Likins, PE, Brent R. Robinson, PE, and 

Matthew L. Becker, EI.

8. PERFORMING ORGANIZATION 

REPORT NO.

9. PERFORMING ORGANIZATION NAME AND ADDRESS

Ryan R. Berg & Associates, Inc.

2190 Leyland Alcove

Woodbury, MN 55125

10. WORK UNIT NO.

11. CONTRACT OR GRANT NO.

DTFH61-11-D-00049

12. SPONSORING AGENCY NAME AND ADDRESS

National Highway Institute

U.S. Department of Transportation

Federal Highway Administration, Washington, DC 20590

13. TYPE OF REPORT & PERIOD 

COVERED

Final Report

14. SPONSORING AGENCY CODE

15. SUPPLEMENTARY NOTES

FHWA COTR: Heather Shelsta

FHWA Technical Working Group: Naser Abu-Hejlah, PhD, PE; Scott Anderson, PE; and

Silas Nichols, PE

16. ABSTRACT

This document presents information on the analysis, design, and construction of driven 

pile foundations for highway structures. This document updates and replaces FHWA 

NHI-05-042 and FHWA NHI-05-043 as the primary FHWA guidance and reference 

document on driven pile foundations. The manual addresses design aspects including 

subsurface exploration, laboratory testing, pile selection, aspects of geotechnical and 

structural limit states, as well as technical specifications. Construction aspects including 

static load tests, dynamic tests, rapid load tests, wave equation analyses, dynamic 

formulas and development of driving criteria, as well as pile driving equipment, pile 

driving accessories, and monitoring of pile installation inspection are also covered. Step 

by step procedures are included for most analysis procedures and design examples.

17. KEY WORDS

Driven pile foundations, foundation economics, site 

characterization, geomaterial properties, axial compression 

resistance, axial tension resistance, lateral resistance, pile 

groups, specifications, nominal resistance determination

tests, construction monitoring and quality assurance. 

18. DISTRIBUTION STATEMENT

No restrictions.

19. SECURITY CLASSIF.

Unclassified

20. SECURITY CLASSIF.

Unclassified

21. NO. OF PAGES

541 

22. PRICE
CONVERSION FACTORS

Approximate Conversions to SI Units Approximate Conversions from SI Units

When You Know Multiply By To Find When You Know Multiply By To Find

(a) Length

inch (in.) 25.4 millimeter (mm) millimeter (mm) 0.039 inch (in.)

foot (ft) 0.305 meter (m) meter (m) 3.28 foot (ft)

yard (yd) 0.914 meter (m) meter (m) 1.09 yard (yd)

mile (mi) 1.61 kilometer (km) kilometer (km) 0.621 mile (mi)

(b) Area

square inches (in2) 645.2 square millimeters (mm2) square millimeters (mm2) 0.0016 square inches (in2) 

square feet (ft2) 0.093 square meters (m2) square meters (m2) 10.764 square feet (ft2) 

Acres (ac) 0.405 hectares (ha) hectares (ha) 2.47 Acres (ac)

square miles (mi2) 2.59 square kilometers (km2) square kilometers (km2) 0.386 square miles (mi2) 

square inches (in2) 645.2 square millimeters (mm2) square millimeters (mm2) 0.0016 square inches (in2) 

(c) Volume

fluid ounces (oz) 29.57 milliliters (mL) milliliters (mL) 0.034 fluid ounces (oz)

Gallons (gal) 3.785 liters (L) liters (L) 0.264 Gallons (gal)

cubic feet (ft3) 0.028 cubic meters (m3) cubic meters (m3) 35.32 cubic feet (ft3) 

cubic yards (yd3) 0.765 cubic meters (m3) cubic meters (m3) 1.308 cubic yards (yd3) 

(d) Mass

ounces (oz) 28.35 grams (g) grams (g) 0.035 ounces

pounds (lb) 0.454 kilograms (kg) kilograms (kg) 2.205 pounds

short tons (2000 lb) (T) 0.907 megagrams (tonne) (Mg) megagrams (tonne) (Mg) 1.102 short tons (2000 lb)

(e) Force

pound (lb) 4.448 Newton (N) Newton (N) 0.2248 pound (lb)

(f) Pressure, Stress, Modulus of Elasticity

pounds per square foot (psf) 47.88 Pascals (Pa) Pascals (Pa) 0.021 pounds per square foot (psf)

pounds per square inch (psi) 6.895 kiloPascals (kPa) kiloPascals (kPa) 0.145 pounds per square inch (psi)

(g) Density

pounds per cubic foot (pcf) 16.019 kilograms per cubic meter 

(kgm3) 

kilograms per cubic meter 

(kgm3) 

0.0

624 pounds per cubic feet (pcf)

(h) Temperature

Fahrenheit temperature (oF) 5/9(oF- 32) Celsius temperature (oC) Celsius temperature (oC) 9/5(oC)+ 32 Fahrenheit temperature (oF)

Notes: 

A compliant version can be found here: https://www.fhwa.dot.gov/publications/convtabl.cfm

1) The primary metric (SI) units used in civil engineering are meter (m), kilogram (kg), second (s), Newton (N), and Pascal (Pa=N/m2). 

2) In a "soft" conversion, an English measurement is mathematically converted to its exact metric equivalent.

3) In a "hard" conversion, a new rounded metric number is created that is convenient to work with and remember.
PREFACE

The purpose of this manual is to provide updated, state-of-the-practice information 

for the design and construction of driven pile foundations in accordance with the 

Load and Resistance Factor Design (LRFD) platform. Engineers and contractors 

have been designing and installing pile foundations for many years. During the past 

three decades, the industry has experienced several major improvements including 

newer and more accurate methods of predicting and measuring geotechnical 

resistance, vast improvements in design software, highly specialized and 

sophisticated equipment for pile driving, and improved methods of construction 

control. Previous editions of the FHWA Design and Construction of Driven Pile 

Foundations manual were published 1985, 1996, and 2006 and chronical the many 

changes in design and construction practice over the past 30 years. This two 

volume edition, GEC-12, serves as the FHWA reference document for highway 

projects involving driven pile foundations. 

Volume I, FHWA-NHI-16-009, covers the foundation selection process, site 

characterization, geotechnical design parameters and reporting, selection of pile 

type, geotechnical aspects of limit state design, and structural aspects of limits state 

design. Volume II, FHWA-NHI-16-010, addresses static load tests, dynamic testing 

and signal matching, rapid load testing, wave equation analysis, dynamic formulas, 

contract documents, pile driving equipment, pile accessories, driving criteria, and 

construction monitoring. Comprehensive design examples are presented in 

publication FHWA-NHI-16-064. 

Throughout this manual, numerous references will be made to the names of 

software or technology that are proprietary to a specific manufacturer or vendor. 

Please note that the FHWA does not endorse or approve commercially available 

products, and is very sensitive to the perceptions of endorsement or preferred 

approval of commercially available products used in transportation applications. Our 

goal with this development is to provide recommended technical guidance for the 

safe design and construction of driven pile foundations that reflects the current state 

of practice and provides information on advances and innovations in the industry. 

To accomplish this, it is necessary to illustrate methods and procedures for design 

and construction of driven pile foundations. Where proprietary products are 

described in text or figures, it is only for this purpose.
The primary audience for this document is: agency and consulting engineers 

specialized in geotechnical and structural design of highway structures; engineering 

geologists and consulting engineers providing technical reviews, or who are 

engaged in the design, procurement, and construction of driven pile foundations 

This document is also intended for management, specification and contracting 

specialists, as well as for construction engineers interested in design and contracting 

aspects of driven pile systems.

This document draws material from the three earlier FHWA publications in this field; 

FHWA-DP-66-1 by Vanikar (1985), FHWA HI 97-013 and FHWA HI 97-014 by 

Hannigan et al. (1998), and FHWA NHI-05-042 and FHWA NHI-05-043 by Hannigan 

et al. (2006). Photographs without specific acknowledgement in this two volume 

document are from these previous editions, their associated training courses, or 

from the consulting practice of GRL Engineers, Inc.

The following individuals were part of the Ryan R. Berg & Associates internal peer 

review team and are acknowledged for their technical advice and contributions to 

this version of the document: 

Mr. Jerry DiMaggio – Applied Research Associates, Inc.

Mr. Van Komurka – Wagner Komurka Geotechnical Group, Inc.

Mr. Billy Camp – S&ME, Inc.

Dr. Brian Anderson – Auburn University
i 

TABLE OF CONTENTS

LIST OF TABLES ...................................................................................... ix

LIST OF FIGURES .................................................................................... xi

9 STATIC LOAD TESTING .................................................................. 1

9.1 GENERAL............................................................................... 1

9.1.1 Reasons and Prerequisites for a Static Load Test 

Program....................................................................... 1

9.1.2 Developing a Static Load Test Program ......................... 2

9.1.3 When and Where to Load Test ...................................... 3

9.1.4 Effective Use of Load Test Information........................... 5

9.1.4.1 Design Stage.................................................. 5

9.1.4.2 Construction Stage ......................................... 5

9.1.4.3 Load Test Databases ...................................... 6

9.1.5 Resistance Factors for Static Load Testing..................... 7

9.2 AXIAL COMPRESSION LOAD TEST........................................ 8

9.2.1 Compression Load Test Equipment ............................... 9

9.2.2 Recommended Axial Compression Test Loading 

Method ......................................................................14

9.2.3 Presentation and Interpretation of Axial Compression 

Test Results ...............................................................16

9.2.4 Limitations of Axial Compression Tests ........................19

9.3 TENSION LOAD TEST ..........................................................20

9.3.1 Tension Load Test Equipment .....................................20

9.3.2 Tension Test Loading Methods....................................21

9.3.3 Presentation and Interpretation of Tension Test 

Results ......................................................................23

9.4 LATERAL LOAD TEST ..........................................................24

9.4.1 Lateral Load Test Equipment.......................................24

9.4.2 Lateral Test Loading Methods .....................................28

9.4.3 Presentation and Interpretation of Lateral Test 

Results ......................................................................29

9.5 LOAD TRANSFER EVALUATIONS.........................................32

9.5.1 Use of Telltales...........................................................32

9.5.2 Use of Strain Gages....................................................34

9.5.3 Determination of Residual Load ...................................41
ii

9.6 PRACTICAL ISSUES AND CONSIDERATIONS ......................46

9.7 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS .........48

REFERENCES.........................................................................................50

10 DYNAMIC TESTING AND SIGNAL MATCHING ANALYSIS..............53

10.1 BACKGROUND.....................................................................53

10.2 APPLICATIONS FOR DYNAMIC TESTING METHODS ...........54

10.2.1 Nominal Resistance ....................................................54

10.2.2 Hammer and Driving System Performance ...................55

10.2.3 Driving Stresses and Pile Integrity................................55

10.3 RESISTANCE FACTORS FOR DYNAMIC TESTING ...............56

10.4 DYNAMIC TESTING ..............................................................57

10.5 BASIC WAVE MECHANICS ...................................................63

10.6 DYNAMIC TESTING METHODOLOGY...................................72

10.6.1 Nominal Resistance Determination by Case Method .....73

10.6.2 Soil Resistance Distributions .......................................77

10.6.3 Energy Transfer..........................................................77

10.6.4 Driving Stresses .........................................................79

10.6.5 Pile Integrity ...............................................................81

10.6.6 Signal Matching ..........................................................82

10.7 CONSIDERATIONS IN TEST SPECIFICATION AND 

IMPLEMENTATION...............................................................91

10.8 PRESENTATION AND INTERPRETATION OF DYNAMIC 

TEST RESULTS....................................................................94

10.8.1 Field Results ..............................................................94

10.8.2 Evaluation of Hammer and Drive System 

Performance...............................................................98

10.8.3 Test Record Illustrating Problematic Hammer 

Performance............................................................. 103

10.8.4 Test Record Illustrating Pile Damage ......................... 104

10.8.5 Test Records Illustrating Soil Setup............................ 104

10.8.6 Test Records Illustrating Relaxation ........................... 108

10.8.7 Reporting of Dynamic Test Results ............................ 110

10.9 CASE HISTORY.................................................................. 112

10.10 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS ....... 114

10.11 PRACTICAL ISSUES AND CONSIDERATIONS .................... 116

REFERENCES....................................................................................... 119

11 RAPID LOAD TESTING ................................................................ 123

11.1 REQUIREMENTS FOR RAPID LOAD TESTS ....................... 123

11.2 BACKGROUND ON RAPID LOAD TEST METHODS............. 125
iii

11.2.1 Combustion Gas and Reaction Mass Apparatus 

(Statnamic)............................................................... 125

11.2.2 Cushioned Drop Weight Systems............................... 130

11.3 RAPID LOAD TEST APPLICATIONS.................................... 132

11.4 RAPID LOAD TEST INTERPRETATION METHODS.............. 132

11.4.1 Unloading Point Method (UPM) ................................. 133

11.4.2 Modified Unloading Point Method (MUP) .................... 138

11.4.3 Segmental Unloading Point Method (SUP) ................. 138

11.4.4 Fully Mobilized UPM ................................................. 139

11.4.5 Sheffield Method for Cohesive Soils........................... 140

11.4.6 Loading Rate Reduction Factors................................ 141

11.4.7 Resistance Factors ................................................... 143

11.5 LATERAL LOADING APPLICATION..................................... 144

11.6 CASE HISTORY.................................................................. 145

11.7 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS ....... 146

11.8 PRACTICAL ISSUES AND CONSIDERATIONS .................... 147

REFERENCES....................................................................................... 149

12 WAVE EQUATION ANALYSIS ...................................................... 153

12.1 WAVE EQUATION ANALYSIS INTRODUCTION................... 153

12.2 WAVE PROPAGATION........................................................ 154

12.3 WAVE EQUATION METHODOLOGY ................................... 155

12.4 WAVE EQUATION APPLICATIONS ..................................... 159

12.5 WAVE EQUATION EXAMPLES............................................ 162

12.5.1 Example 1 – General Bearing Graph.......................... 162

12.5.2 Example 2 – Constant Capacity / Variable Stroke 

Option...................................................................... 167

12.5.3 Example 3 – Drivability Studies.................................. 167

12.5.4 Example 4 – Tension and Compression Stress 

Control..................................................................... 174

12.5.5 Example 5 – Use of Soil Setup .................................. 179

12.5.6 Example 6 – Driving System Characteristics............... 182

12.5.7 Example 7 – Assessment of Pile Damage .................. 187

12.5.8 Example 8 – Selection of Wall Thickness ................... 193

12.5.9 Example 9 – Evaluation of Vibratory Driving ............... 197

12.6 ANALYSIS DECISIONS FOR WAVE EQUATION 

MODELING ......................................................................... 203

12.6.1 Selecting the Proper Approach .................................. 203

12.6.2 External Combustion Hammer Consideration ............. 205

12.6.3 Diesel Hammer Considerations ................................. 206
iv

12.6.4 Vibratory Hammer Considerations ............................. 207

12.6.5 Batter Pile Considerations ......................................... 208

12.6.6 Hammer and Pile Cushion Considerations.................. 209

12.6.7 Selection of Soil and Rock Parameters....................... 210

12.6.8 Pile Modeling Considerations .................................... 213

12.6.9 Comparison with Dynamic Measurements .................. 215

12.7 WAVE EQUATION ANALYSIS INPUT .................................. 217

12.7.1 Other Analysis Options.............................................. 237

12.8 WAVE EQUATION ANALYSIS OUTPUT AND 

INTERPRETATION.............................................................. 245

12.9 PLOTTING OF WAVE EQUATION RESULTS ....................... 251

12.10 SUGGESTIONS FOR PROBLEM SOLVING ......................... 251

12.11 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS ....... 255

12.12 PRACTICAL ISSUES AND CONSIDERATIONS .................... 256

REFERENCES....................................................................................... 258

13 DYNAMIC FORMULAS................................................................. 261

13.1 BACKGROUND................................................................... 261

13.1.1 Historical Accuracy of Dynamic Formulas ................... 262

13.1.2 Basic Limitations with Dynamic Formulas ................... 265

13.2 RESISTANCE FACTORS FOR DYNAMIC FORMULAS......... 266

13.3 DYNAMIC FORMULAS........................................................ 268

13.3.1 FHWA Modified Gates Formula ................................. 268

13.3.2 AASHTO Modified Engineering News Formula ........... 269

13.3.3 Other Dynamic Formulas........................................... 269

13.3.3.1 Washington State DOT Pile Driving Formula. 269

13.3.3.2 Minnesota Pile Formula 2012 (MPF12)......... 271

13.4 DYNAMIC FORMULA CASE HISTORY ................................ 273

13.5 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS ....... 277

13.6 PRACTICAL ISSUES AND CONSIDERATIONS .................... 277

REFERENCES....................................................................................... 279

14 CONTRACT DOCUMENTS........................................................... 281

14.1 OVERVIEW OF PLAN AND SPECIFICATION 

REQUIREMENTS................................................................ 281

14.2 GENERIC DRIVEN PILE SPECIFICATION ........................... 284

14.2.1 SECTION X.01 DESCRIPTION ................................ 286

14.2.2 SECTION X.02 SUBMITTALS AND APPROVALS ..... 287

14.2.3 SECTION X.03 MATERIALS .................................... 295

14.2.4 SECTION X.04 DRIVING EQUIPMENT AND 

APPURTENANCES.................................................. 296
v 

14.2.5 SECTION X.05 DETERMINATION OF NOMINAL 

RESISTANCE .......................................................... 303

14.2.6 SECTION X.06 PREPARATION AND DRIVING ........ 316

14.2.7 SECTION X.07 METHOD OF MEASUREMENT ........ 325

14.2.8 SECTION X.08 BASIS OF PAYMENT....................... 329

REFERENCES....................................................................................... 331

15 PILE DRIVING EQUIPMENT......................................................... 333

15.1 CRANES............................................................................. 333

15.2 DEDICATED AND UNIVERSAL RIGS................................... 336

15.3 LEADS................................................................................ 338

15.4 TEMPLATES....................................................................... 348

15.5 HELMETS ........................................................................... 349

15.6 HAMMER CUSHIONS ......................................................... 352

15.7 PILE CUSHIONS ................................................................. 353

15.8 HAMMERS.......................................................................... 355

15.8.1 Hammer Energy Concepts ........................................ 358

15.9 DROP HAMMERS ............................................................... 359

15.10 SINGLE ACTING AIR/STEAM HAMMERS ............................ 361

15.11 DOUBLE ACTING AIR/STEAM HAMMERS........................... 363

15.12 DIFFERENTIAL ACTING AIR/STEAM HAMMERS................. 366

15.13 SINGLE ACTING (OPEN END) DIESEL HAMMERS.............. 368

15.14 DOUBLE ACTING (CLOSED END) DIESEL HAMMERS........ 371

15.15 SINGLE ACTING HYDRAULIC HAMMERS........................... 373

15.16 DOUBLE ACTING HYDRAULIC HAMMERS ......................... 375

15.17 VIBRATORY HAMMERS...................................................... 376

15.18 RESONANT HAMMERS ...................................................... 378

15.19 IMPACT HAMMER SIZE SELECTION .................................. 379

15.20 HAMMER KINETIC ENERGY MONITORING ........................ 380

15.21 NOISE SUPPRESION EQUIPMENT..................................... 384

15.22 UNDERWATER NOISE SUPPRESSION EQUIPMENT.......... 385

15.23 FOLLOWERS...................................................................... 386

15.24 JETTING............................................................................. 388

15.25 PREDRILLING..................................................................... 389

15.26 SPUDDING ......................................................................... 390

15.27 EQUIPMENT SUBMITTALS ................................................. 390

REFERENCES....................................................................................... 393

16 PILE ACCESSORIES ................................................................... 395

16.1 TIMBER PILES.................................................................... 395

16.1.1 Pile Toe Attachments ................................................ 395
vi

16.1.2 Attachments at Pile Head .......................................... 397

16.1.3 Splices ..................................................................... 397

16.2 STEEL H-PILES .................................................................. 398

16.2.1 Pile Toe Attachments ................................................ 398

16.2.2 Splices ..................................................................... 400

16.3 STEEL PIPE PILES ............................................................. 403

16.3.1 Pile Toe Attachments ................................................ 403

16.3.2 Splices ..................................................................... 408

16.3.3 Constrictor Plates ..................................................... 413

16.4 PRECAST CONCRETE PILES ............................................. 414

16.4.1 Pile Toe Attachment.................................................. 414

16.4.2 Splices ..................................................................... 416

16.5 WELDED SPLICES AND TOE ATTACHMENTS.................... 422

16.5.1 Welding Surface Preparation..................................... 423

16.5.2 Temperature Requirements During Welding and 

Splicing .................................................................... 426

16.5.3 Welding Pile Toe Accessories.................................... 427

16.5.4 Welded Splice Checklist............................................ 429

16.5.4.1 Preparation................................................. 429

16.5.4.2 During Welding ........................................... 429

16.5.4.3 Final Weld Inspection.................................. 430

REFERENCES....................................................................................... 431

17 DRIVING CRITERIA ..................................................................... 433

17.1 DEVELOPMENT OF THE PILE DRIVING CRITERIA............. 433

17.2 PRACTICAL AND ABSOLUTE REFUSAL ............................. 435

17.3 PRACTICAL ISSUES AND CONSIDERATIONS .................... 436

17.4 EXAMPLES FOR ESTABLISHING DRIVING CRITERIA ........ 441

17.4.1 Driving Criterion – Example 1 .................................... 442

17.4.2 Driving Criterion – Example 2 .................................... 443

REFERENCES....................................................................................... 445

18 CONSTRUCTION MONITORING OF PILE INSTALLATION............ 447

18.1 MONITORING NEEDS BASED ON THE PROJECT 

DELIVERY METHOD........................................................... 448

18.2 ITEMS TO BE MONITORED ................................................ 448

18.3 REVIEW OF PROJECT PLANS AND SPECIFICATIONS ....... 451

18.4 INSPECTORS TOOLS......................................................... 451

18.5 INSPECTION OF PILES PRIOR TO AND DURING 

INSTALLATION................................................................... 452

18.5.1 Timber Piles ............................................................. 452
vii

18.5.2 Precast Concrete Piles.............................................. 453

18.5.3 Steel H-Piles ............................................................ 455

18.5.4 Steel Pipe Piles ........................................................ 455

18.6 INSPECTION OF DRIVING EQUIPMENT ............................. 456

18.7 INSPECTION OF DRIVING EQUIPMENT DURING 

INSTALLATION................................................................... 461

18.7.1 Drop Hammers ......................................................... 462

18.7.2 Single Acting Air/Steam Hammers ............................. 463

18.7.3 Double Acting or Differential Air/ Steam Hammers ...... 468

18.7.4 Single Acting Diesel Hammers................................... 471

18.7.5 Double Acting Diesel Hammers ................................. 476

18.7.6 Single Acting Hydraulic Hammers .............................. 482

18.7.7 Double Acting Hydraulic Hammers............................. 485

18.7.8 Resonant Hammers .................................................. 489

18.7.9 Vibratory Hammers................................................... 489

18.8 INSPECTION OF TEST OR INDICATOR PILES.................... 490

18.9 INSPECTION OF PRODUCTION PILES ............................... 493

18.10 DRIVING RECORDS AND REPORTS .................................. 502

18.11 SAFETY.............................................................................. 506

REFERENCES....................................................................................... 507

A. LIST OF FHWA/ NHI RESOURCES RELEVANT TO DEEP 

FOUNDATIONS ........................................................................... 509

B. LIST OF ASTM AND AASHTO PILE DESIGN AND TESTING 

SPECIFICATIONS........................................................................ 515

C. PILE HAMMER INFORMATION .................................................... 519
viii
ix

LIST OF TABLES

Table 9-1 Resistance Factors Based on Static Load Testing of Driven 

Piles (modified from AASHTO 2014)........................................ 7

Table 9-2 Standard Loading Schedule (after ASTM D3966) ...................29

Table 10-1 Summary of Case Damping Factors for RSP and RMX 

Equations ............................................................................75

Table 10-2 Pile Damage Guidelines (after Rausche and Goble 1979).......81

Table 10-3 Description of Typical Dynamic Test Output Codes ................96

Table 12-1 Bearing Graph Numerical Results ....................................... 164

Table 12-2 Example 6: Proposed Hammer and Driving Systems ............ 183

Table 12-3 Example 7 – Maxima of Forces, Stresses and Other 

Variables for Required Nominal Resistance of 380 kips ........ 190

Table 12-4 Example 7 – Maxima of Forces, Stresses and Other 

Variables for at Refusal Driving Condition (Nominal 

Resistance of 500 kips) ...................................................... 190

Table 12-5 Example 8 – Wave Equation Results for Four Potential Pipe 

Pile Wall Thicknesses at 2000 kip Nominal Resistance ......... 195

Table 12-6 Soil Information for Vibratory Sheet Pile Driving in Example 

9 ....................................................................................... 199

Table 12-7 Nominal Resistance and Wave Equation Calculated Results 

for Vibratory Drivability Analysis. Total Drive Time – 4 

Minutes ............................................................................. 202

Table 12-8 Soil Parameters in ST Analysis for Granular Soil Types ........ 229

Table 12-9 Soil Parameters in ST Analysis for Cohesive Soil Types ....... 229

Table 12-10 GRLWEAP Extrema Output ................................................ 248

Table 12-11 Suggested Use of the Wave Equation to Solve Field 

Problems ........................................................................... 252

Table 12-12 Wave Equation Analysis Problems ...................................... 254

Table 13-1 Mean Values and Coefficients of Variation for Various 

Methods ............................................................................ 265

Table 13-2 Resistance Factors for Dynamic Formulas (modified from 

AASHTO 2014).................................................................. 267

Table 13-3 Pile Types and Sizes Contained in WSDOT Formula 

Database........................................................................... 271

Table 13-4 Case History – Comparison of Calculated Nominal 

Resistances ....................................................................... 275
x 

Table 13-5 Case History – Comparison of Factored Resistance ............. 276

Table 13-6 INDOT Dynamic Formula and Dynamic Testing Comparison. 278

Table 14-1 Items to Include in a Driven Pile Specification ...................... 283

Table 14-2 Resistance Factors for Field Determination Methods (after 

AASHTO 2014).................................................................. 293

Table 14-3 Alternate Approval Method Minimum Pile Hammer 

Requirements .................................................................... 294

Table 14-4 Common Time Delay for Restrike Based Upon Soil Type...... 308

Table 14-5 Basis of Payment ............................................................... 330

Table 15-1 Typical Pile Hammer Characteristics and Uses .................... 356

Table 15-2 Preliminary Hammer Energy Requirements.......................... 380

Table 16-1 Summary of Precast Concrete Pile Splices (after Mullins 

and Sen 2015) ................................................................... 416

Table 17-1 Example 1 Summary of Requirements and Results .............. 442

Table 18-1 Overview of Key Construction Monitoring and Inspection 

Items and Agency Inspection Duties on Design-Bid-Build 

and Design Build Contracts................................................. 449

Table 18-2 Common Problems and Indicators for Air/Steam Hammers 

(after Williams Earth Sciences 1995) ................................... 466

Table 18-3 Common Problems and Indicators for Single Acting Diesel 

Hammers (after Williams Earth Sciences 1995) ................... 476

Table 18-4 Common Problems and Indicators for Double Acting Diesel 

Hammers (after Williams Earth Sciences 1995) ................... 480

Table 18-5 Common Problems and Indicators for Single Acting 

Hydraulic Hammers (after Williams Earth Sciences 1995) ..... 483

Table 18-6 Common Problems and Indicators for Double Acting 

Hydraulic Hammers............................................................ 487

Table 18-7 Common Pile Installation Problems and Possible Solutions... 498
xi

LIST OF FIGURES

Figure 9-1 Axial compression static load test diagram............................... 9

Figure 9-2 Static load test setup diagram. ..............................................11

Figure 9-3 Load test application and monitoring system. .........................12

Figure 9-4 Load test movement monitoring components. ........................13

Figure 9-5 MnDOT reaction beam and load test setup (courtesy 

MnDOT). .............................................................................13

Figure 9-6 4000-Ton mobile load test frame (courtesy Caltrans). .............15

Figure 9-7 Static load test setup using weighted platform. .......................15

Figure 9-8 Typical load-movement curve for axial compression load 

test......................................................................................17

Figure 9-9 Tension test on pipe pile using reaction piles (courtesy of 

Besix)..................................................................................22

Figure 9-10 Tension test on H-pile using mats and cribbing (courtesy of 

WKG2).................................................................................22

Figure 9-11 Typical tension load test load-movement curve. .....................23

Figure 9-12 Lateral load test on 14 inch and 16 inch O.D. concrete filled 

pipe piles (courtesy of WKG2)................................................26

Figure 9-13 Jack for lateral load test (courtesy of WKG2). .........................26

Figure 9-14 Spherical bearing plate and load cell for lateral load test 

(courtesy of WKG2)...............................................................27

Figure 9-15 Multiple inclinometer string components for lateral load test

(courtesy of WKG2

)...............................................................27

Figure 9-16 Typical Shape Accel Array (SSA) (courtesy of 

Measurand). ........................................................................28

Figure 9-17 Typical lateral load test, pile head load-deflection curve. .........30

Figure 9-18 Plot of lateral load test measured deflected shape versus 

depth...................................................................................31

Figure 9-19 Comparison of measured and COM624P predicted load 

deflection behavior versus depth (after Kyfor et al. 1992). .......31

Figure 9-20 Diagram of telltale rods installed on pile (modified from 

Kyfor et al. 1992).................................................................. 33

Figure 9-21 Vibrating wire strain gage sister bars for concrete 

embedment..........................................................................35

Figure 9-22 Vibrating wire strain gage with welded anchor blocks and 

protective channel. ...............................................................35
xii

Figure 9-23 Waterproof electrical resistance strain gage bolted on pipe 

pile with protective channel (courtesy of Besix).......................36

Figure 9-24 Electrical resistance strain gage on sister bars in concrete 

pile casting bed. ...................................................................36

Figure 9-25 Multiple externally mounted strain gages (2 on each web 

face) located in soil resistance free area during static load 

test (courtesy WKG2)............................................................38

Figure 9-26 Tangent modulus method for determining elastic modulus 

(modified from Fellenius 2001). .............................................40

Figure 9-27 Plot of axial load versus strain gage elevation for load 

transfer assessment. ............................................................41

Figure 9-28 Example of residual load effects on load transfer 

evaluation. ...........................................................................42

Figure 9-29 Mobilized shaft resistance hysteresis loop for pile during 

static load test (after Fellenius 2014). ....................................43

Figure 9-30 Mobilized toe resistance for pile during static load test. 

(after Fellenius 2014)............................................................44

Figure 9-31 Instrumented static load test results and soil resistance 

distribution (modified from Fellenius 2002b with data from 

Altaee et al. 1992)................................................................45

Figure 9-32 Variation between applied pile head load determined from 

load cell and jack pressure gage (after Fellenius 1984)...........47

Figure 9-33 Load cell response based on jack and bearing plate size 

(after Geokon 2013). ............................................................48

Figure 10-1 Pile preparation for dynamic testing. ......................................58

Figure 10-2 a) Accelerometer, strain gage, and WiFi transmitter 

(courtesy Pile Dynamics) and b) combined strain gage and 

accelerometer (courtesy Allnamics). ......................................59

Figure 10-3 Embedded strain gage and accelerometer unit being cast 

into concrete pile (courtesy Radise International)....................60

Figure 10-4 Embedded resistance strain gage and accelerometer 

mounted on sister bar being cast into concrete pile 

(courtesy of Pile Dynamics, Inc.). ..........................................60

Figure 10-5 Pile Driving Analyzer processing unit (courtesy of Pile 

Dynamics, Inc.). ...................................................................62

Figure 10-6 Embedded Data Collector system (courtesy of Radise 

International)........................................................................62

Figure 10-7 Free end wave mechanics. ...................................................64

Figure 10-8 Force and velocity times (EA/C) records versus time for 

free end...............................................................................65
xiii

Figure 10-9 Fixed end wave mechanics. ..................................................67

Figure 10-10 Force and velocity times (EA/C) records versus time for 

fixed end..............................................................................68

Figure 10-11 Soil resistance effects on force and velocity records (after 

Hannigan 1990). ..................................................................69

Figure 10-12 Typical force and velocity records for various soil 

resistance conditions. ...........................................................71

Figure 10-13 Examples of standard Case Method, RSP, and maximum 

Case Method, RMX, calculations for estimates of nominal 

resistance. ........................................................................... 76

Figure 10-14 Energy transfer computation. ................................................78

Figure 10-15 Example compression and tension stress calculation..............80

Figure 10-16 Schematic of CAPWAP signal matching method. ...................84

Figure 10-17 Factors most influencing signal matching analysis. .................85

Figure 10-18 Example of signal matching iteration matching process...........86

Figure 10-19 Example of signal matching graphical output..........................87

Figure 10-20 Example signal matching output “Summary Results” table. .....89

Figure 10-21 Example signal matching output “Extrema” table. ...................90

Figure 10-22 Example signal matching output “Case Method” table.............90

Figure 10-23 Example signal matching output “Pile Profile and Pile 

Model” table.........................................................................91

Figure 10-24 APPLE drop weight system...................................................93

Figure 10-25 Typical dynamic test display from Pile Driving Analyzer. .........94

Figure 10-26 Energy transfer ratios for select hammer and pile 

combinations........................................................................99

Figure 10-27 Histograms of energy transfer ratio for diesel hammers on 

(a) steel piles and (b) concrete/timber piles. ......................... 100

Figure 10-28 Histograms of energy transfer ratio for single acting 

air/steam hammers on (a) steel piles and (b) 

concrete/timber piles. ......................................................... 101

Figure 10-29 Histograms of energy transfer ratio for single acting 

hydraulic hammers on (a) steel piles and (b) 

concrete/timber piles. ......................................................... 102

Figure 10-30 Example dynamic test records on pre-igniting diesel 

hammer............................................................................. 103

Figure 10-31 Example dynamic test records indicating pile damage. ......... 105

Figure 10-32 Photographs of extracted pile showing damage.................... 105

Figure 10-33 Example test record illustrating soil setup - end of initial 

driving data........................................................................ 106
xiv

Figure 10-34 Example test record illustrating soil setup - beginning of 

restrike data....................................................................... 106

Figure 10-35 Example test record illustrating relaxation - end of initial 

driving data........................................................................ 109

Figure 10-36 Example test record illustrating relaxation - beginning of 

restrike data....................................................................... 109

Figure 10-37 Typical tabular presentation of dynamic test results. ............. 111

Figure 10-38 Typical graphical presentation of dynamic test results 

versus depth. ..................................................................... 112

Figure 11-1 Typical axial compressive force pulse (after ASTM D7383). .. 124

Figure 11-2 Statnamic concept (courtesy of Berminghammer 

Foundation Equipment)....................................................... 126

Figure 11-3 1000 kip hydraulic catch device on prestressed concrete 

pile (courtesy of Applied Foundation Testing). ...................... 127

Figure 11-4 Statnamic test in progress with gravel catch mechanism - 

9,000 kip device (courtesy of Applied Foundation Testing). ... 128

Figure 11-5 Statnamic raw force and displacement measurements 

versus time (courtesy of Berminghammer Foundation 

Equipment). ....................................................................... 129

Figure 11-6 Statnamic load versus displacement (courtesy of 

Berminghammer Foundation Equipment)............................. 129

Figure 11-7 Drop weight rapid load test systems: (a) Fundex system 

(courtesy Foundation Constructors Inc.) and (b) 

Hybridnamic system. .......................................................... 131

Figure 11-8 APPLE 32 ton modular rapid load test system with 

transducer. ........................................................................ 131

Figure 11-9 Free body diagram of pile forces in a Statnamic test (after 

Middendorp et al. 1992)...................................................... 135

Figure 11-10 Five stages of a Statnamic test (after Middendorp et al. 

1992)................................................................................. 135

Figure 11-11 Derived Statnamic load displacement curve with rate 

effects (courtesy of Berminghammer Foundation 

Equipment). ....................................................................... 138

Figure 11-12 Normalized resistance versus normalized displacement 

(after Miyasaka et al. 2009)................................................. 140

Figure 11-13 Variation of the UPM loading rate reduction factor versus 

the soil liquid limit (after Brown and Powell 2013). ................ 142

Figure 11-14 Lateral Statnamic test on nine pile group (courtesy of Utah 

State University)................................................................. 144

Figure 11-15 Statnamic test result (courtesy Minnesota DOT)................... 145
xv

Figure 12-1 Wave propagation in a pile (adapted from Cheney and 

Chassie 2000).................................................................... 155

Figure 12-2 Typical wave equation pile and soil models. ......................... 157

Figure 12-3 Example 1 – Problem profile. .............................................. 163

Figure 12-4 Typical wave equation bearing graph. Top plot – driving 

stresses; bottom plot – nominal resistance and diesel 

hammer stroke................................................................... 165

Figure 12-5 Constant capacity analysis (inspector’s chart). ..................... 168

Figure 12-6 Example 3 – Problem profile. .............................................. 169

Figure 12-7 Nominal resistance, calculated blow count and stresses for 

before and after densification and nominal resistance for 

the after scour condition...................................................... 173

Figure 12-8 Drivability results for H-pile with corresponding concrete 

pile results. ........................................................................ 173

Figure 12-9 Example 4 – Problem profile. .............................................. 174

Figure 12-10 Bearing graphs for easy driving condition, two pile cushion 

thicknesses. ....................................................................... 176

Figure 12-11 Bearing graphs for early low energy and end of driving high 

energy driving condition. ..................................................... 178

Figure 12-12 Example 5 –- Problem profile. ............................................. 180

Figure 12-13 Using a bearing graph with soil setup................................... 181

Figure 12-14 Example 6 – Problem profile. .............................................. 183

Figure 12-15 Bearing graph – for two hammers with equivalent potential 

energy and large toe quake................................................. 184

Figure 12-16 Bearing graph – for two hammers with equivalent potential 

energy and normal toe quake.............................................. 186

Figure 12-17 Example 7 –- Problem profile. ............................................. 187

Figure 12-18 Wave equation bearing graph for proposed driving system.... 189

Figure 12-19 Wave equation bearing graphs for damaged and 

undamaged pile. ................................................................ 192

Figure 12-20 Example 8 – Problem profile. .............................................. 194

Figure 12-21 Bearing graphs for open end pipe piles with wall 

thicknesses of 0.50, 0.75, 1.00, and 1.25 inches................... 196

Figure 12-22 Example 9 – Problem profile. .............................................. 198

Figure 12-23 Nominal resistance and wave equation calculated 

penetration time for vibratory drivability analysis. .................. 201

Figure 12-24 Pile driving and equipment data form................................... 218

Figure 12-25 GRLWEAP help window for main input form......................... 220

Figure 12-26 Job information window. ..................................................... 221

Figure 12-27 Select hammer window....................................................... 221
xvi

Figure 12-28 Analysis type window. ........................................................ 222

Figure 12-29 Pile input window. .............................................................. 223

Figure 12-30 Area calculator window....................................................... 225

Figure 12-31 Hammer cushion window.................................................... 226

Figure 12-32 (a) Soil profile input window for soil type based static soil 

analysis and (b) Soil parameter input window for bearing 

graph analysis.................................................................... 228

Figure 12-33 Window for entering up to 10 nominal resistance values for 

bearing graph and inspector’s chart analyses....................... 230

Figure 12-34 Resistance gain/loss factors in main input form for 

drivability analysis. ............................................................. 232

Figure 12-35 S1 form for soil resistance vs. depth input. ........................... 233

Figure 12-36 New profile window for the SA static analysis. ...................... 235

Figure 12-37 Soil property input window for the SA static analysis............. 236

Figure 12-38 Depths, modifiers input form. .............................................. 237

Figure 12-39 Completed main input for a simple bearing graph. ................ 238

Figure 12-40 Batter/Inclination input window. ........................................... 239

Figure 12-41 Slack/Splice information input window. ................................ 240

Figure 12-42 Data entry screen for non-uniform piles................................ 241

Figure 12-43 Numeric options window. .................................................... 242

Figure 12-44 Damping options window. ................................................... 243

Figure 12-45 Stroke options window for diesel hammers........................... 244

Figure 12-46 Output options window. ...................................................... 246

Figure 12-47 Hammer model, hammer options, and driving system 

output................................................................................ 247

Figure 12-48 Pile, soil, and analysis options............................................. 249

Figure 12-49 Extrema table output. ......................................................... 250

Figure 12-50 Final summary for bearing graph analysis. ........................... 250

Figure 13-1 Log normal probability density function for four resistance 

predictions (after Rausche et al. 1996)................................. 264

Figure 14-1 Drive system submittal form. ............................................... 289

Figure 15-1 Crawler crane pile driving rig with lattice boom and 

swinging leads. .................................................................. 334

Figure 15-2 Truck crane pile driving rig with lattice boom and swinging 

leads. ................................................................................ 335

Figure 15-3 Rough terrain pile driving rig with telescoping boom and 

offshore leads. ................................................................... 336

Figure 15-4 Dedicated pile driving system.............................................. 337

Figure 15-5 Universal rig setup for pile driving (courtesy Bauer-Pileco). ... 338

Figure 15-6 Typical lead types (after DFI Publication 1981)..................... 339
xvii

Figure 15-7 Swinging lead systems (after DFI Publication 1981).............. 340

Figure 15-8 Stabbing guide at bottom of box leads. ................................ 342

Figure 15-9 Pile gate with latch for use on truss lead (courtesy 

Berminghammer). .............................................................. 342

Figure 15-10 Fixed lead systems (after DFI Publication 1981). .................. 343

Figure 15-11 Fixed lead system with air compressor mounted as 

counterweight (courtesy FHWA Demo 66). .......................... 344

Figure 15-12 Vertical travel lead system (courtesy Berminghammer)......... 345

Figure 15-13 Lead configuration for batter piles (after DFI publication 

1981)................................................................................. 346

Figure 15-14 Typical offshore lead configuration. ..................................... 347

Figure 15-15 Offshore lead and pile guide (left) with pipe pile helmet 

(right) (courtesy Berminghammer). ...................................... 347

Figure 15-16 Template system................................................................ 348

Figure 15-17 Template elevation effects on batter piles (after Passe 

1994)................................................................................. 349

Figure 15-18 Helmet components (after DFI Publication 1981).................. 350

Figure 15-19 Cross section of cracked helmet with striker plate and 

aluminum and micarta hammer cushion (courtesy of 

WKG2)............................................................................... 350

Figure 15-20 One piece helmets for (a) concrete pile and (b) pipe pile. ...... 351

Figure 15-21 Inserts for (a) H-pile and (b) varying pipe pile diameter and 

wall thickness..................................................................... 351

Figure 15-22 Typical hammer cushion materials....................................... 353

Figure 15-23 Plywood pile cushions: (left) used and (right) new................. 354

Figure 15-24 Pile hammer classification................................................... 355

Figure 15-25 Typical drop hammer.......................................................... 360

Figure 15-26 Schematic of a single acting air/steam hammer.................... 362

Figure 15-27 (a) Single acting air, and (b) differential acting air 

hammers. .......................................................................... 363

Figure 15-28 Schematic of a double acting air/steam hammer................... 365

Figure 15-29 Schematic of a differential acting air/steam hammer. ............ 367

Figure 15-30 Schematic of a single acting diesel hammer......................... 369

Figure 15-31 (a) Single acting diesel and, (b) double acting diesel 

hammers. .......................................................................... 371

Figure 15-32 Schematic of a double acting diesel hammer........................ 372

Figure 15-33 Schematic of hydraulic hammers: (a) single acting and (b) 

double acting. .................................................................... 374

Figure 15-34 (a) Single acting hydraulic, and (b) double acting hydraulic 

hammers. .......................................................................... 375
xviii

Figure 15-35 Schematic of a vibratory hammer. ....................................... 377

Figure 15-36 Vibratory hammer installing an H-pile (courtesy ICE). ........... 377

Figure 15-37 Resonant hammer installing an H-pile (courtesy of 

Resonance Technology). .................................................... 379

Figure 15-38 IHC hydraulic hammer kinetic energy readout panel. ............ 381

Figure 15-39 Proximity switch attachment for Berminghammer diesel 

hammer (courtesy Berminghammer).................................... 382

Figure 15-40 Proximity switches, readout device and driving log trigger 

switch for Berminghammer diesel hammer (courtesy 

Berminghammer). .............................................................. 382

Figure 15-41 Proximity switches and transmitter on retrofitted diesel 

hammer............................................................................. 383

Figure 15-42 E-Saximeter wireless kinetic energy readout device.............. 383

Figure 15-43 Noise shroud for IHC hydraulic hammer............................... 384

Figure 15-44 Bubble ring with containment device (courtesy of WSDOT)... 386

Figure 15-45 (a) Follower and (b) follower in use drving H-piles................. 387

Figure 15-46 Dual jet system mounted on concrete pile (courtesy of 

Florida DOT)...................................................................... 388

Figure 15-47 Jet/Punch system (courtesy of Florida DOT). ....................... 389

Figure 15-48 Solid flight auger predrilling system. .................................... 390

Figure 15-49 (a) Pipe pile and (b) H-pile spuds. ....................................... 391

Figure 15-50 Drive system submittal form. ............................................... 392

Figure 16-1 Timber pile toe attachments: (a) Pile boot and (b) Pile point 

(courtesy Skyline Steel). ..................................................... 396

Figure 16-2 Timber pile with pile boot attached....................................... 396

Figure 16-3 Banded timber pile head. .................................................... 397

Figure 16-4 Damaged H-pile driven without pile toe protection 

(courtesy Wisconsin DOT). ................................................. 399

Figure 16-5 Styles of H-pile driving shoes (courtesy Skyline Steel). ......... 399

Figure 16-6 Pile shoe welded to pile toe using fillet welds along exterior 

flanges. ............................................................................. 400

Figure 16-7 Full penetration groove weld with backer bars. ..................... 401

Figure 16-8 H-pile splicer...................................................................... 401

Figure 16-9 Partially completed splice using H-pile splicer. ..................... 402

Figure 16-10 Closure plate typical of closed end steel pipe pile. ................ 404

Figure 16-11 Conical toe attachments for closed end pipe piles: (a) 

Rounded (b) Pointed conical tip (courtesy Skyline Steel)...... 405

Figure 16-12 Conical toe attachment welded on a closed end steel pipe 

pile. ................................................................................... 405
xix

Figure 16-13 Rock crusher driving shoe welded on a closed end steel 

pipe pile............................................................................. 406

Figure 16-14 Cutting shoes for open end pipe piles: (a) Inside flange and 

(b) Outside flange (courtesy Skyline Steel)........................... 407

Figure 16-15 Inside flange cutting shoe welded to open end steel pipe 

pile. ................................................................................... 407

Figure 16-16 Outside flange cutting shoe on open end steel pipe pile. ....... 408

Figure 16-17 Splices for pipe piles........................................................... 409

Figure 16-18 Backing rings for pipe pile splicing (a) with pins (courtesy 

Skyline Steel) and (b) without.............................................. 409

Figure 16-19 Backing ring tack welded to bottom pile. .............................. 410

Figure 16-20 Welded pipe pile sections. .................................................. 411

Figure 16-21 Drive-fit mechanical friction splicer for pipe pile (courtesy 

Skyline Steel)..................................................................... 412

Figure 16-22 Drive-fit mechanical friction splicer on pipe pile..................... 412

Figure 16-23 Constrictor plate installed inside pipe pile............................. 413

Figure 16-24 Pile toe attachments on concrete pile: (a) Flat shoe 

(courtesy Titus Steel) and (b) Oslo point. ............................. 414

Figure 16-25 Oslo point for concrete pile.................................................. 415

Figure 16-26 Steel H-pile cast into toe of prestressed concrete pile. .......... 415

Figure 16-27 Commonly used prestressed concrete pile splices (after 

PCI 1993). ......................................................................... 417

Figure 16-28 Cement/Epoxy-dowel splice (after Bruce and Herbert 

1974)................................................................................. 418

Figure 16-29 Epoxy-dowel splice: (a) pile with core holes and (b) splice 

in progress......................................................................... 419

Figure 16-30 Welded ICP splice: (a) beveled steel fitting and (b) welded 

splice................................................................................. 419

Figure 16-31 Kie-Lock splice................................................................... 420

Figure 16-32 Post-tensioned splice (a) splice joint and (b) post 

tensioning strands in upper section (courtesy University of 

South Florida). ................................................................... 420

Figure 16-33 Sleeve splices: (a) Hawaiian can and (b) Bruns splice. ......... 421

Figure 16-34 Cracked weld on pipe pile splice.......................................... 422

Figure 16-35 Broken weld on H-pile splice. .............................................. 423

Figure 16-36 Beveled edges on pipe pile in preparation for weld. .............. 424

Figure 16-37 Grinding edges of web and flange in preparation for weld. .... 424

Figure 16-38 H-pile splice preparation without adequate root opening or 

bevel. ................................................................................ 425

Figure 16-39 Misalignment between pile sections and poor quality weld. ... 425
xx

Figure 16-40 Extracted H-pile upper section. Weld cracked during 

driving and extracted section retrieved without H-pile 

splicer or lower pile section. ................................................ 426

Figure 16-41 Heating pile splice location prior to welding. ......................... 427

Figure 16-42 Unsatisfactory spot welds used for shoe attachment............. 428

Figure 16-43 Full groove weld between pile and toe attachment................ 428

Figure 18-1 Key components of the pile installation inspection process.... 450

Figure 18-2 Key components of the pile installation process (modified 

from Williams Earth Science 1995). ..................................... 452

Figure 18-3 Check of Blue Nylon hammer cushion material before use. ... 458

Figure 18-4 Damaged aluminum and micarta hammer cushion. .............. 458

Figure 18-5 New pile cushion being inserted into helmet......................... 459

Figure 18-6 Helmet with new pile cushion installed. ................................ 460

Figure 18-7 Air compressor display panel. ............................................. 464

Figure 18-8 Inspector’s form for single and differential acting air/steam 

hammers. .......................................................................... 467

Figure 18-9 Inspector’s form for enclosed double acting air/steam 

hammers. .......................................................................... 470

Figure 18-10 Fuel Pumps: (a) fixed four step pump and (b) variable fuel 

pump................................................................................. 473

Figure 18-11 Hydraulic pump for fuel pump adjustments........................... 473

Figure 18-12 Inspector’s form for single acting diesel hammers................. 475

Figure 18-13 Typical external bounce chamber pressure gauge. ............... 478

Figure 18-14 Inspector’s form for double acting diesel hammers. .............. 481

Figure 18-15 Inspector’s form for single acting hydraulic hammers. ........... 484

Figure 18-16 IHC hydraulic hammer read-out panel................................. 486

Figure 18-17 Inspector’s form for hydraulic hammers................................ 488

Figure 18-18 Driving sequence of displacement pile groups (after Passe 

1994)................................................................................. 496

Figure 18-19 Pile driving log. .................................................................. 504

Figure 18-20 Daily inspection report. ....................................................... 505
xxi

LIST OF SYMBOLS

A - Pile cross sectional area (9.2) (9.5) (10.4) (10.5) (10.6) (12.8);

Slope of tangent modulus line (9.5). 

Ac - Cross sectional area of concrete (9.5).

a - Acceleration of pile (11.4).

a(t) - Acceleration measured at the gage location (10.4).

B - Y-intercept of the tangent modulus line (9.5). 

b - Pile width or diameter (9.2); 

bpm - Blow per minute (15.13) (18.7).

C - Wave speed of pile material (10.5) (10.6) (11.3) (12.8). 

C4 - Damping constant for Statnamic test (11.4).

D - Pile width for WEAP quake calculation(12.5);

Dw - Wave length (11.3).

E - Elastic modulus of pile material ( 9.2) (9.5) (10.4) (10.5) (10.6) (12.8). 

Ec - Elastic modulus of concrete (9.5).

Ed - Developed hammer energy (13.3). 

Ep - Energy computed at the gage location (10.6). 

Esm - Secant modulus (9.5). 

Est - Elastic modulus of steel (9.5). 

F - Axial force in plane of gage (9.5); Force (10.5).

F(t) - Force computed at gage location (10.4) (10.6).

Fa - Statnamic inertia load (11.4). 

Feff - Hammer efficiency factor (13.3).

Fmax - Maximum measured force applied during testing (11.4).

Fp - Statnamic pore water pressure force (11.4). 

Fstn - Statnamic induced load (11.4).

Ft - Force measured on pile head (11.4).

Fu - Statnamic static soil resistance force (11.4).

Fv - Statnamic dynamic soil resistance force (11.4). 

f’c - Concrete compressive strength at 28 days (14.2).

fpe - Effective prestress in concrete (14.2).

g - Acceleration due to gravity (15.8).

h - Ram stroke (13.1) (15.8) (15.13) (18.7). 

J - Dimensionless damping factor based on soil type (10.6).

KE - Kinetic Energy (15.8).

L - Total pile length (10.5) (10.6) (11.3); 
xxii

Pile length below dial gages or gage location (9.2) (10.6); 

Length of pile between two measuring points under no load condition 

(9.5).

Mt - Tangent modulus (9.5).

m - Mass of pile (11.4); Ram mass (15.8).

Nb - Pile penetration resistance with units of (blows/inch) (13.3).

Nft - Pile penetration resistance with units of (blows/foot) (13.3).

Nt - Toe bearing capacity coefficient (12.7). 

Nw - Wave number (11.3).

Q - Test load (9.2) (9.5).

Qavg - Average load in the pile between two points (9.5).

Qf - Failure load (9.2) (9.3). 

RSP - Case method static nominal resistance (10.6).

RTL - Total static and dynamic resistance on pile during driving (10.6).

Rn - Nominal resistance (11.4) (13.1)

Rndr - Nominal driving resistance (13.3).

Rp - Nominal toe resistance (9.5).

R1 - Deflection reading at upper measurement location (9.5).

R2 - Deflection reading at lower measurement location (9.5).

sb - Set per blow (13.1); Permanent pile set (13.3).

sf - Measured pile head movement at failure (9.2) (9.3).

tL - Load duration (11.4).

tumax - Time of maximum displacement for Statnamic test (11.4).

t1 - Time of initial impact (10.6).

t2 - Time of reflection of initial impact from pile toe (t1+2L/C) (10.6).

t4 - Time at beginning of Stage 4 for Statnamic test (11.4).

V - Velocity of pile (11.4). 

V(t) - Velocity measured at gage location (10.4). 

Vi - Impact velocity (15.8). 

Vo - Normalizing constant (11.4).

Vstatic - Velocity used to determine parameters αα and ββ for Sheffield Method 

(11.4).

W - Ram weight (13.1). 

WD(t) - Downward traveling wave, wave down (10.6).

WDm - Wave down, measured (10.6).

WU(t) - Upward traveling wave, wave up (10.6).

WUc - Wave up, computed (10.6).

WUm - Wave up, measured (10.6).

α - Sheffield Method rate parameter (11.4). 

β - Sheffield Method rate parameter (11.4). 
xxiii

ξ - Loading rate reduction factor (11.4).

Δ - Elastic deformation of pile (9.2).

Δε - Change of strain from one load increment to the next (9.5).

Δσ - Change of stress from one load increment to the next (9.5).

ε - Strain measured in gage (9.5).

ε(t) - Measured strain at time, t (10.4).

σ - Normal stress (pressure) on plane of failure, stress (9.5).

φdyn - Resistance factor (based on the construction control method) (13.2).
xxiv
xxv

LIST OF ACRONYMNS

AASHTO - American Association of State Highway and Transportation Officials

ASTM - American Society for Testing and Materials

BL - Blast load

BOR - Beginning of Restrike 

BR - Vehicular braking force

CD - Consolidated Drained triaxial test

CE - Vehicular centrifugal force

CED - Closed End Diesel hammer

CEP - Closed End Pipe

COR - Coefficient of Restitution

CPT - Cone Penetration Test

CPTu - Piezo Cone Penetration Test

CR - Force effects due to creep 

CT - Vehicular collision force

CU - Consolidated Undrained triaxial test

CV - Vessel collision force

DA - Design Angular Distortion

DC - Dead load components and attachments

DD - Downdrag 

DMT - Dilatometer test

DW - Wearing surface and utilities

DWT - Deadweight tonnage 

EH - Horizontal earth pressure 

EL - Locked-in stress 

EOD - End of Drive

EQ - Earthquake load

ER - SPT hammer efficiency as determined by energy measurements

ES - Earth surcharge 

EV - Vertical earth pressure

FHWA - Federal Highway Administration

FR - Friction load

I.D. - Inner diameter

IC - Ice load

IM - Vehicular dynamic load allowance

KE - Kinetic Energy
xxvi

LL - Liquid Limit; Vehicular Live Load

LS - Live Load Surcharge

LVDT - Linear Variable Differential Transformer

MUP - Modified Unloading Point Method

NHI - National Highway Institute

O.D. - Outer Diameter

OED - Open Ended Diesel hammer

OEP - Open Ended Pipe

PE - Potential Energy

PGA - Peak Ground Acceleration coefficient

PI - Plasticity Index

PL - Plastic Limit; Pedestrian Live Load

PS - Secondary forces from post-tensioning

RSA - Residual Stress Analysis

SE - Force effect due to settlement

SH - Force effects due to shrinkage

SPT - Standard Penetration Test

SRD - Soil Resistance to Driving

SUP - Segmental Unloading Point Method 

TG - Force effect due to temperature gradient

TU - Force effect due to uniform temperature

UPM - Unloading Point Method

UU - Unconsolidated Undrained triaxial test

VST - Vane shear test

WA - Water load and steam pressure

WD - Downward traveling wave, Wave Down

WEAP - Wave Equation Analysis Program

WL - Wind on live load

WS - Wind load on structure

WU - Upward traveling wave, Wave Up
1

CHAPTER 9

STATIC LOAD TESTING

9.1 GENERAL

Static load testing of piles is the most accurate method of determining load capacity. 

Depending upon the size of the project and other project variables, static load tests 

may be performed either during the design stage or construction stage. 

Conventional load test types include the axial compression, axial tension and lateral 

load tests. 

The purpose of this chapter is to provide an overview of static testing and its 

importance as well as to describe the basic test methods and interpretation 

techniques. For additional details on static load testing, reference should be made 

to publications listed at the end of this chapter including the most recent ASTM test 

standards as well as the NYSDOT Static Pile Load Test Manual, GCP-18 (2007).

9.1.1 Reasons and Prerequisites for a Static Load Test Program

Static load testing provides the engineer with valuable design verification information 

on the nominal resistance and deformation response of test or production piles. 

Reasons to perform static load tests include the following:

1. Provide information for design verification or design refinement as well as 

information for construction verification or construction procedure modification 

based on measured load-deflection or measured load-transfer data.

2. Confirm the ability of the subsurface materials to provide the nominal 

geotechnical resistance.

3. Determine or calibrate resistance factors for new static design methods, local 

or regional geologic conditions, or dynamic or rapid load test analysis 

methods or procedures.

4. Determine p-y response of laterally loaded piles.
2

Static load test program development and oversight should be performed by an 

experienced foundation engineer. Static load tests are frequently used to establish a 

comparison baseline for design, testing, and construction techniques. The quality of 

static load tests is always important, but should be of utmost importance when 

performing research or updating current LRFD resistance factors. Poorly performed 

tests or analyses can result in unsafe or grossly uneconomical foundations.

In order to adequately plan and implement a static load testing program, the 

following information should first be obtained or developed.

1. A detailed subsurface exploration program at the test location. A load test is 

not a substitute for a subsurface exploration program.

2. Well defined subsurface stratigraphy including engineering parameters of 

geomaterials and identification of groundwater conditions.

3. Static analyses to economically select appropriate pile type(s) and length(s) 

as well as to select appropriate location(s) for load test(s).

9.1.2 Developing a Static Load Test Program

The goal and objectives of a static load test program should be clearly established, 

while the type and frequency of tests should be selected to provide the required 

knowledge for final design purposes or for construction verification. A significantly 

different level of effort and instrumentation is required if the goal of the load test 

program is simply to confirm the nominal geotechnical resistance or if detailed loadtransfer information is desired for final design or improvement of design efficiency. 

The following items should be considered during the test program planning so that 

the program provides the desired information.

1. The capacity of the loading apparatus (reaction system and jack) should be 

specified so that the test pile(s) may be loaded to the geotechnical nominal 

resistance. A loading apparatus designed to load a pile to only to the 

anticipated nominal resistance is usually insufficient to obtain geotechnical 

failure. Hence, the true nominal geotechnical resistance cannot be 

determined, and the full benefit from performing a static load test to improve 

design efficiency is not realized. The jack capacity should be between 120% 

and 150% of the anticipated maximum load to be applied.
3

2. Specifications should require the use of a load cell and spherical bearing 

plate as well as linear variable displacement transducers (LVDT’s) or dial 

gages with sufficient travel to allow accurate measurements of load and 

movement at the pile head. LVDT’s or dial gages with insufficient travel that 

must be shimmed and reset during the test should be prohibited. Where 

possible, deformation measurements should also be made at the pile toe and 

at intermediate points to allow for an evaluation of shaft and toe resistances. 

The load cell capacity should be between 120% and 150% of the anticipated 

maximum load to be applied.

3. The load test program should be supervised by a foundation engineer 

experienced in this field of work.

4. A test pile installation record should be maintained with installation details 

appropriately noted. Too often, only the hammer model and driving 

resistance are recorded on a test pile log. Additional items such as hammer 

stroke (particularly at final driving), fuel/energy setting, hammer cushion 

materials and dimensions, pile cushion material and cushion thickness when 

new and replaced, accurately determined final set, details on any installation 

aids used such as predrilling and their depths, start and ending driving times, 

stopping depths and associated time durations for splicing, equipment 

maintenance, etc., should be recorded.

5. Use of dynamic monitoring on the load test pile is recommended for estimates 

of the nominal resistance at the time of driving including the soil resistance 

distribution, evaluation of drive system performance, calculation of driving 

stresses, and refinement of wave equation input parameters. 

9.1.3 When and Where to Load Test

Static load testing during either the design or construction phase should be 

performed to achieve a desired goal. This usually involves determining the nominal 

geotechnical resistance or load-transfer information at a representative site location. 

The number of load tests needed and their location should be selected based on the 

site variability. The AASHTO resistance factor for determination of the nominal 

resistance by static load testing is based on performing one load test per site 

condition. Site variability is discussed in Section 5.5.3 and in AASHTO (2014).

The following criteria, adapted from FHWA-SA-91-042 by Kyfor et al. (1992), 

summarize conditions when static pile load testing can be effectively utilized:
4

1. When substantial cost savings can be realized. This is often the case on 

large projects involving either friction piles (to determine that estimated pile 

lengths can be reduced) or end bearing piles (to determine that a higher 

nominal resistance can be used).

2. When the nominal geotechnical resistance is uncertain due to limitations of an 

engineer's experience base, or due to unusual site or project conditions.

3. When subsurface conditions vary considerably across the project, but can be 

delineated into zones of similar conditions. Static tests can then be 

performed in representative areas to delineate foundation variation.

4. When a significantly higher nominal geotechnical resistance is contemplated 

relative to typical practice.

5. When significant time dependent changes in nominal resistance are 

anticipated as a result of soil setup or relaxation. 

6. When a reliable assessment of axial tension resistance or lateral deflection is 

important.

7. Verification of new design or testing methods. 

8. When new pile types, large diameter open end pipe piles, and/or pile 

installation procedures are utilized.

9. When existing piles will be reused to support a new structure with heavier 

loads.

10. When, during construction, the estimated nominal resistance using dynamic 

formulas or dynamic analysis methods differs significantly from the estimated 

resistance at that depth determined by static analysis. For example, H-piles 

that "run" when driven into loose to medium dense sands and gravels.

11. When developing LRFD calibration based on local and regional geologic 

conditions.
5

9.1.4 Effective Use of Load Test Information

9.1.4.1 Design Stage

The best information for design of a pile foundation is provided by the results of a 

load testing program conducted during the design phase. The number of static 

tests, types of piles to be tested, method of driving and test load requirements 

should be selected by the geotechnical and structural engineers responsible for 

design. A cooperative effort between the two disciplines is necessary. The following 

are the advantages of load testing during the design stage.

a. Allows load testing of several different pile types, pile sections, and lengths 

resulting in the design selection of the most economical pile foundation.

b. Confirm drivability to minimum penetration requirements and suitability of 

nominal geotechnical resistance at the estimated pile penetration depths.

c. Establishes preliminary driving criteria for production piles. 

d. The availability of pile driving information to construction project bidders 

should reduce their bid "contingency."

e. Reduces potential for contract disputes related to pile driving problems.

f. Allows the results of load test program to be reflected in the final design and 

specifications.

9.1.4.2 Construction Stage

Load testing at the start of construction may be the only practical time for testing on 

some projects that cannot justify the cost of a design stage program. Construction 

stage static load tests are invaluable to confirm that the design loads are 

appropriate, to establish the final design lengths for production piles, and to assure 

that the pile installation procedure is satisfactory. Perhaps most importantly, these 

results can be used to refine the estimated pile lengths shown on the plans and 

establish minimum pile penetration requirements.
6

9.1.4.3 Load Test Databases

As mentioned in Paikowsky (2004) and Abu-Hejleh (2015), load test databases are 

needed for design method reliability calibrations and to improve the geotechnical 

design of production foundations. High quality, complete, load test databases offer 

the ability to compare design and construction verification methods with full scale 

load test results. The information may allow the increase of the resistance factor for 

a specific design method, or provide designers with preferable foundation 

alternatives and construction methods given similar site conditions. 

One such study was performed in 2010 by the Oregon Department of Transportation 

(DOT), where they evaluated driven pile resistance factors based upon wave 

equation analyses (Smith et al. 2011). At that time, the AASHTO recommended 

resistance factor, φdyn, for wave equation analysis was 0.40. This resistance factor 

resulted in increased pile foundation costs within the LRFD framework. Therefore, 

the Oregon DOT performed a calibration study to assess the reliability of a higher 

resistance factor for wave equation analysis. After reviewing several databases, the 

study recommended a wave equation analysis resistance factor of 0.55 for initial 

driving and 0.40 for restrikes. As a result of this research and other efforts, the 

recommended resistance factor for wave equation analysis in the AASHTO (2010) 

design specifications was changed from φdyn =0.40 to φdyn =0.50. 

The FHWA Deep Foundation Load Test Database (DFLTD) contains data from 

several state highway agencies, FHWA offices and international sources. Although 

it is no longer updated, the DFLTD is available upon request by contacting the 

FHWA and contains approximately 1300 load test results from 1985 to 2003. 

Additional load test databases have been developed by several state transportation 

agencies. 

It is recommended that load test results and accompanying geotechnical and pile 

installation information be stored in databases. The respective report section of the 

ASTM standard for axial compression, axial tension, and lateral load tests identifies 

the complete site, installation, and load test details that should maintained in a 

database. Maintaining this information allows the database to be effective for 

development and calibration of new design or construction control methods.
7 

9.1.5 Resistance Factors for Static Load Testing

Resistance factors applicable to the nominal axial geotechnical resistance 

determined from a static load test are provided in Table 9-1. The resistance factor 

varies depending on whether the load test is in axial compression or tension as well 

as whether dynamic testing is performed. A resistance factor is not provided for 

lateral load testing as this is generally controlled by deformation criteria in the 

service limit state rather than by the strength limit state.

Table 9-1 Resistance Factors Based on Static Load Testing of Driven Piles

(modified from AASHTO 2014)

Condition Resistance Determination Method Resistance Factor

Nominal

Axial Compression 

Resistance of Single 

Pile, φdyn

Driving criteria established by 

successful static load test of at 

least one pile per site condition and 

dynamic testing of at least two piles 

per site condition, but no less than 

2% of the production piles.

0.80

Nominal 

Axial Compression 

Resistance of Single 

Pile, φdyn

Driving criteria established by 

successful static load test of at 

least one pile per site condition 

without dynamic testing.

0.75

Nominal 

Axial Tension 

Resistance of Single 

Pile, φup

Static load test. 0.60
8

9.2 AXIAL COMPRESSION LOAD TEST

Piles are most frequently tested in axial compression. However, they can also be 

tested in axial tension as well as for their lateral resistance. Figure 9-1 illustrates the 

basic mechanism of performing an axial compression load test. This mechanism 

normally includes the following steps:

1. The pile is loaded incrementally from the pile head using some predetermined 

loading sequence, or it can be loaded at a continuous, constant rate.

2. During the test, readings are recorded of the time, applied load, and pile head 

movement. Strain measurements at other points along the pile shaft can be 

obtained concurrently to yield load transfer details and unit shaft resistances.

3. A load movement curve is plotted.

4. The geotechnical nominal resistance and the movement at the nominal 

resistance are determined by one of the several methods of interpretation.

5. Movement is usually measured only at the pile head. However, the pile can 

also be instrumented with telltales to determine movement anywhere along 

the pile. Telltales (solid rods protected by tubes) are shown in Figure 9-1. 

6. Telltales can be used to determine the strain between telltale locations. 

Strain gages, attached to, or embedded within the pile can be used to

determine the strain at discrete locations along the pile length. Both of these 

strain measurements can be used to estimate load transfer along the pile 

shaft.
9 

Figure 9-1 Axial compression static load test diagram.

9.2.2 Compression Load Test Equipment

ASTM D1143-07 (re-approved 2013) recommends several alternative systems for 

(1) applying compression load to the pile, and (2) measuring movements. Most 

often, compression loads are applied by hydraulically jacking against a beam that is 

anchored by piles or ground anchors, or by jacking against a weighted platform. The 

primary means of measuring the load applied to the pile should be with a calibrated 

load cell. The jack pressure from a calibrated pressure gage should also be 

recorded and be used as a secondary means of calculating the load applied to the 

pile. To minimize eccentricities in the applied load, a spherical bearing plate should 

be included in the load application arrangement.
10

Axial pile head movements are usually measured by LVDT's or dial gages that 

measure movement between the pile head and independently supported reference 

beam. The ASTM standard requires dial gages or LVDT's to have a minimum of 2 

inches of travel and a precision of at least 0.01 inches. However, it is preferable and 

highly recommended to have gages with a minimum travel of 4 inches particularly for 

long piles with large elastic deformations under load and with a precision of 0.001 

inches. A minimum of two dial gages or LVDT's mounted equidistant from the pile 

head, equidistant from the center of the pile, and diametrically opposite should be 

used. In many instances, four dial gages or LVDT’s mounted in diametrically 

opposite pairs may be advantageous.

A redundant backup system consisting of a scale, mirror, and wire system should be 

provided with a scale precision of 0.01 inches. The backup system should also be 

mounted on two diametrically opposite pile faces. Both the reference beams and 

backup wire system are to be independently supported with a clear distance of at 

least 5 times the pile diameter and not less than 8 feet between their supports and 

the test pile or reaction piles. A remote backup system consisting of a survey level 

should also be used in case reference beams or wire systems are disturbed during 

the test. The survey level can also monitor reaction pile movement where reaction 

piles are used.

ASTM D1143 specifies that the clear distance between a test pile and reaction piles 

be at least 5 times the maximum diameter of the reaction pile or test pile (whichever 

has the greater diameter if not the same pile type) but not less than 8 feet. If a 

weighted platform is used, the clear distance between cribbing supporting the 

weighted platform and the test pile should exceed 5 feet.

A schematic of a typical compression load test setup is presented in Figure 9-2. 

Photographs of the typical load application and movement monitoring components 

are presented in Figures 9-3 and 9-4. To improve the accuracy of the load cell 

readings, the loading arrangement shown in Figure 9-3 should include steel bearing 

plates between the hydraulic jack ram and the load cell, and between the load cell 

and spherical bearing plates. These bearing plates are included in the loading 

arrangement shown in Figure 9-5, and are discussed in greater detail in Section 9.6. 

The MnDOT mobile load frame attached to reaction piles is shown in Figure 9-5 with 

the load test being setup enclosed in a temporary heated structure for winter 

weather control. Wood reference beams are used to minimize temperature 

fluctuations on the reference beams.
11

Figure 9-2 Static load test setup diagram.
12

Figure 9-3 Load test application and monitoring system.
13

Figure 9-4 Load test movement monitoring components.

Figure 9-5 MnDOT reaction beam and load test setup (courtesy MnDOT).
14

A typical compression load test arrangement using reaction piles is presented in 

Figure 9-6 and a weighted platform arrangement is shown in Figure 9-7. It is 

recommended that load and movement measurements be recorded remotely using 

electronic instrumentation so that personnel are a significant distance from the load 

application / reaction system. Additional details on load application as well as pile 

head load and movement measurements may be found in ASTM D1143.

9.2.3 Recommended Axial Compression Test Loading Method

It is essential that standardized load testing procedures be followed. Several loading 

procedures are detailed in ASTM D1143, Standard Test Method for Deep 

Foundations Under Static Axial Compressive Load. The quick load test procedure is 

recommended by AASHTO (2014). This procedure is recommended because the 

load test can typically be performed in 2 to 4 hours thereby reducing construction 

delays and load test costs. However, alternative test procedures are also provided 

in ASTM D1143. 

In the quick test procedure, the load is applied in increments of 5% of the anticipated 

nominal geotechnical resistance. This loading increment should produce on the 

order of 20 data points being recorded before reaching the geotechnical nominal 

resistance load thus allowing a detailed load-movement curve to be plotted. 

Load increments should be held no more than 15 minutes and no less than 4 

minutes, using the same time interval for all load increments. Readings of load and 

gross movement are to be recorded at 0.5, 1, 2, and 4 minutes after the load 

increment increase, with 8 and 15 minute readings if longer load increments have 

been selected. This procedure is to continue until the geotechnical nominal 

resistance or jack capacity is reached within the safe structural design limitations of 

the pile and reaction system. ASTM D1143 states that a longer hold time may be 

considered at the geotechnical nominal resistance to access creep. Note that this is 

not feasible if plunging movement under the applied load has occurred. If plunging 

movement occurs, the pump pressure should be locked off and the system should 

be allowed to come to equilibrium under the maximum resisted load. Upon reaching 

and holding the maximum load or reaching the load where plunging movement 

occurred, the pile is unloaded in 5 to 10 equal decrements which are each held for 

no more than 15 minutes and no less than 4 minutes. Readings of load and gross 

movement are to be recorded at 0.5, 1, 2, and 4 minutes (as well as 8 and 15 

minutes) after each load reduction, including the zero load, which can be held longer 

to assess rebound.
15

Figure 9-6 4000-Ton mobile load test frame (courtesy Caltrans).

Figure 9-7 Static load test setup using weighted platform.
16

9.2.4 Presentation and Interpretation of Axial Compression Test Results

The results of axial compression load tests should be presented in a report 

conforming to the requirements of ASTM D1143. A load-movement curve similar to 

the one shown in Figure 9-8 should be plotted for interpretation of load test results. 

The literature abounds with different methods of defining the nominal geotechnical 

resistance from static load tests. Methods of interpretation based on maximum 

allowable gross movements which do not take into account the elastic deformation 

of the pile are not recommended. These methods overestimate the nominal 

resistance of short piles and underestimate the nominal resistance of long piles. 

Methods which account for elastic deformation and are based on failure criterion 

provide a better understanding of pile performance and provide more accurate

results. 

AASHTO (2014) design specifications Article 10.7.3.8.2 provides three interpretation 

criteria of axial compression test results based on pile size. All of the methods 

include consideration of the elastic deformation of the pile. For uniform presentation 

and interpretation of the axial compression load test results, the load and movement 

scales are selected so that the line representing the elastic deformation, Δ, of the 

pile is inclined at an angle of about 20° from the load axis.

The elastic deformation, Δ, for a pile of uniform cross section is computed from:

AE

QL =∆ Eq. 9-1 

Where:

Δ = elastic deformation of pile (inches).

Q = test load (kips).

L = pile length below dial gage or LVDT measurement location (inches).

A = pile cross sectional area (in2

).

E = elastic modulus of pile material (ksi).

AASHTO (2014) axial compression load test interpretation methods are based on an 

offset limit method as proposed by Davisson (1972). An offset limit line parallel to 

the elastic deformation line is plotted as shown in Figure 9-8. The point at which the 

observed load movement curve intersects the offset limit line is by definition the 

nominal geotechnical resistance or failure load. Note that this is a limiting 

deformation based criteria. In Figure 9-8, geotechnical plunging failure was not 
17

achieved, and additional geotechnical resistance develops beyond the defined 

failure load with additional pile head movement. If the load-movement curve does 

not intersect the offset limit line, the pile has a nominal geotechnical resistance in 

excess of the maximum applied test load. 

Figure 9-8 Typical load-movement curve for axial compression load test.

For piles 24 inches or less in diameter, the nominal geotechnical resistance, or 

failure load Qf , is that load which produces a movement of the pile head equal to:

 Eq. 9-2 

Where:

sf = measured pile head movement at the nominal resistance (inches).

Δ = elastic deformation of pile per Equation 9-1 (inches).

b = pile width or diameter (inches).
18

For piles larger than 36 inches in diameter, additional pile toe movement is 

necessary to develop the toe resistance. For these larger diameter piles, the 

nominal geotechnical resistance, or failure load, can be defined as the load which 

produces at movement at the pile head equal to:

bs 30/ f +∆= Eq. 9-3 

Where: 

sf = measured pile head movement at the nominal resistance (inches).

Δ = elastic deformation of pile per Equation 9-1 (inches).

b = pile width or diameter (inches).

For piles greater than 24 inches but less than 36 inches in diameter, linear 

interpolation should be performed between Eq. 9-2 and Eq. 9-3. 

The pile diameter, b, used in Eq. 9-2 and Eq. 9-3 is defined by AASHTO as the pile 

diameter or width of side for square piles. However, for H-piles the value for b is not 

defined. For consistency with AASHTO treatment of square piles, it is 

recommended that the flange width be chosen for b.

It should be noted that the Davisson method is applicable for load tests in which the 

increment of load is held for not more than 1 hour. It is not applicable to load tests 

with long load holds and will underestimate the nominal geotechnical resistance in 

those cases due to the additional pile head movement that occurs during the 

extended load holds.

As discussed in Section 9.2.1, the load applied to the pile head is measured with a 

primary (load cell) and back-up (jack pressure) system. The applied load 

determined from both the primary and secondary system should be in general 

agreement, and should be compared to each other throughout the test to detect any 

major errors or discrepancies. If during initial loading increments, large 

discrepancies are observed between the applied load indicated by the primary and 

secondary systems, the test should be stopped, and the cause of the large 

discrepancy evaluated. Strain gage instrumentation above the ground surface can 

oftentimes quickly determine whether the load measured by the primary system or 

backup system is correct, and therefore which system should be used to control the 

test. If a load test is performed with large unresolved discrepancies, safety issues 

and/or later controversies may arise. If the lower value of applied load is used, and 

the actual applied load is the higher value, a safety issues can occur during the test 

by operating the jack beyond its intended range, or by exceeding the nominal 
19

structural resistance of the test pile or reaction system. If the higher value of applied 

load is used, and the actual applied load is the lower value, the test may be 

terminated prematurely or the nominal resistance determined from the test lower 

than anticipated. This may require the entire load test to be repeated, the use of 

longer pile lengths than necessary, a foundation redesign because of the lower load, 

or revisions to the driving criterion. Therefore whenever large discrepancies occur, 

the applied load should be removed from the test pile (unloaded), the cause of the 

discrepancy identified and corrected, and the load test restarted.

The factored compression resistance is 0.75 to 0.80 of the nominal resistance 

defined by the failure load. This range in the factored compression resistance 

depends on whether the driving criteria is developed solely from the static load test 

results (φdyn = 0.75), or if static load test results and dynamic testing with signal 

matching (φdyn = 0.80) is used.

An axial compression load test will not provide detailed load-transfer information in 

the penetrated geomaterials solely from the pile head load-movement results. 

Additional instrumentation as described in Section 9.5 is required if load-transfer 

information along the pile shaft or the delineation of shaft and toe resistances is 

desired.

9.2.5 Limitations of Axial Compression Tests

Compression load tests can provide a wealth of information for design and 

construction of pile foundations and are the most accurate method of determining 

nominal resistance. However, static load test results cannot be used to account for 

long-term settlement, downdrag from consolidating and settling soils, or to 

adequately represent pile group action. Other shortcomings of static load tests 

include test cost, the time required to setup and complete a test, and the minimal 

information obtained on driving stresses or extent of pile damage (if any). Static 

load test results can also be misleading on projects with highly variable soil 

conditions. If detailed load-transfer assessments are desired, the pile must have 

additional strain gage instrumentation attached or embedded in the pile, which can 

add significant cost.
20

9.3 TENSION LOAD TEST

Tension load tests are performed to determine axial tension (uplift) resistance of 

piles. The uplift resistance is often a significant factor in determining minimum pile 

penetration requirements for pile groups subject to large uplift loading demand 

including cofferdam seals that create large buoyancy forces, cantilever segmental 

bridge construction, as well as seismic, vessel impact, or debris loading. The basic 

mechanics of a tension test are similar to compression load testing, except the pile is 

loaded in tension. Tension load tests are often economical since they can frequently 

be performed by jacking against a reaction beam supported by crane mats rather 

than reaction piles. 

9.3.1 Tension Load Test Equipment

ASTM D3689-07 (re-approved 2013) describes The Standard Method of Testing 

Deep Foundations Under Static Axial Tensile Load. Several alternative systems for 

applying tension load to the pile and measuring movements are provided in this 

standard. Most often, tension loads are applied by centering a hydraulic jack on 

reaction beams supported by reaction piles or by mats and cribbing supported by the 

ground. The jack can be a center hole jack allowing a rod embedded in the pile to 

pass in between the reaction beams and through the jack. This is a common 

arrangement for concrete piles with relatively low tension loads. Alternatively, a 

yoke type of connection can be used to transfer the tension load to the pile. In this 

arrangement, steel rods pass on either side of the reaction beam(s). The rods are 

connected to the pile and to a beam on top of the jack. This is a common 

arrangement for steel piles. The primary means of measuring the load applied to the 

pile should be from a calibrated load cell with the jack load recorded from a 

calibrated pressure gage as backup. A spherical bearing plate should be included in 

the load application arrangement.

Axial pile head movements are usually measured by dial gages or LVDT’s that 

measure movement between the pile head and an independently supported 

reference beam. For tension testing, ASTM requires that the dial gages or LVDT’s 

have a minimum of 2 inches of travel and a precision of at least 0.01 inches. While 

not required, gages with 4 inches of travel are recommended. A minimum of two dial 

gages or LVDT’s mounted equidistant from the point of load application, equidistant 

from the center of the pile, and diametrically opposite should be used. A redundant 

backup system consisting of a scale, mirror, and wire system should be provided 

with a scale precision of 0.01 inches. The backup system should also be mounted 

on two diametrically opposite pile faces. Both the reference beams and backup wire 
21

system are to be independently supported with a clear distance of at least 5 times 

the pile diameter and not less than 8 feet between supports and the test pile or 

reaction piles. A remote backup system consisting of a survey level should also be 

used in case reference beams or wire systems are disturbed during the test. 

Additional details on load application, and pile head load and movement 

measurements may be found in ASTM D3689.

Photographs of tension load test arrangements are presented in Figure 9-9 and 

9-10. In Figure 9-9, a tension load test arrangement for a nearshore situation using 

a reaction beam and reaction piles is presented. In Figure 9-10, a tension load test 

arrangement for a land situation using a reaction beam supported by crane mats and 

timber cribbing is illustrated.

9.3.2 Tension Test Loading Methods

Several loading procedures are detailed in ASTM D3689. The Quick Test method 

where load is applied in increments of 5% of the anticipated nominal geotechnical 

tension resistance is recommended. This loading increment should produce on the 

order of 20 data points being recorded before reaching the nominal geotechnical 

resistance thus allowing a detailed load-movement curve to be plotted. All load 

increments should be held for the same time interval of no more than 15 minutes 

and no less than 4 minutes. 

Readings of load and gross movement are to be recorded at 0.5, 1, 2, and 4 minutes 

after the load increment increase, with 8 and 15 minute readings for longer hold 

times. This procedure is to continue until the nominal geotechnical resistance or 

jack capacity is reached within the safe structural design limitations of the pile and 

reaction system. Upon reaching and holding the maximum load or reaching the load 

where geotechnical pullout failure occurred, the pile is unloaded in 5 to 10 equal 

decrements which are each held for no more than 15 minutes and no less than 4 

minutes. Load and gross movement are to be recorded at 0.5, 1, 2, and 4 minutes 

(as well as 8 and 15 minutes if applicable) after each load reduction, including the 

zero load. Additional optional loading procedures are detailed in ASTM D3689. 

It is generally desirable to test a pile in tension loading to geotechnical pullout failure, 

particularly during a design stage test program. If construction stage tension tests 

are performed on production piles, the piles should be redriven to the original pile 

toe elevation and the previous driving resistance upon completion of the load testing.
22

Figure 9-9 Tension test on pipe pile using reaction piles (courtesy of 

Besix).

Figure 9-10 Tension test on H-pile using mats and cribbing (courtesy of WKG2

).
23

9.3.3 Presentation and Interpretation of Tension Test Results

The results of tension load tests should be presented in a report conforming to the 

requirements of ASTM D3689. A load movement curve similar to the one shown in 

Figure 9-11 should be plotted for interpretation of tension load test results.

A widely accepted method for determining the nominal geotechnical resistance in 

tension loading has not been published. Fuller (1983) reported that acceptance 

criteria for tension tests have included a limit on the gross or net upward movement 

of the pile head, the slope of the load movement curve, or an offset limit method that 

accounts for the elastic lengthening of the pile plus an offset.

Figure 9-11 Typical tension load test load-movement curve.

AASHTO (2014) design specifications recommend tension load tests be evaluated 

using a modified Davisson Method that considers the elastic lengthening of the pile 

plus an offset limit. For tension loading, the suggested offset is 0.15 inches. Hence, 

the load at which the load movement curve intersects the elastic lengthening plus 

0.15 inches is then defined as the nominal tension resistance or failure load. The 

factored tension resistance is then 0.60 of the nominal tension resistance defined by 

the failure load. 
24

9.4 LATERAL LOAD TEST

Lateral load tests are performed on projects where piles are subjected to significant 

lateral loads. The importance of determining pile response to lateral loading has 

greatly increased, particularly with regard to extreme event limit states consideration 

such as seismic and vessel impact. Lateral load testing has also increased due to 

the greater use of noise walls and large overhead signs. Most lateral load tests are 

performed on a single pile with a free head condition. In service, most piles are 

installed in groups having a fixed head condition. Hence, lateral load test results on 

a single pile with a free head pile condition may not be directly applicable to design. 

Accordingly, the primary purpose of lateral load testing is to determine the p-y 

curves modelling soil behavior to be used in the design or to verify the 

appropriateness of the p-y curves on which the design is based. Lateral load tests

are often economical since they can frequently be performed by jacking or pulling 

between two piles. For design stage tests, the two piles can be different pile types 

or pile sections enabling lateral deformation behavior for two pile types to be 

assessed with one lateral load test.

9.4.1 Lateral Load Test Equipment

ASTM D3966-07 (re-approved 2013) describes The Standard Method of Testing 

Deep Foundations Under Lateral Load. Several alternative systems for applying the 

lateral load to the pile and measuring movements are provided in this standard. 

Most often, lateral loads are applied by a hydraulic jack acting against a reaction 

system (e.g. piles, deadman, or weighted platform), or by a hydraulic jack acting 

between two piles. When jacking between two piles, pulling the piles toward one 

another is the preferred arrangement as the system comprises a two-force member 

and so alignment is naturally maintained. However, with this arrangement ASTM 

requires a significant distance between the test piles clear distance of not less than 

20 feet or 20 pile diameters. The primary means of measuring the load applied to 

the pile(s) should be from a calibrated load cell with the jack load recorded from a 

calibrated pressure gage as backup. ASTM D3966 requires a spherical bearing 

plate(s) be included in the load application arrangement unless the load is applied by 

pulling.

Lateral pile head movements are usually measured by dial gages or LVDT's that 

measure movement between the pile head and an independently supported 

reference beam mounted perpendicular to the direction of movement. For lateral 

load testing, ASTM requires the dial gages or LVDT's have a minimum of 3 inches of 

travel and a precision of at least 0.01 inches. For tests on a single pile, one dial 
25

gage or LVDT is mounted on the side of the test pile opposite the point of load 

application. A backup system consisting of a scale, mirror, and wire system should 

be provided with a scale precision of 0.01 inches. The backup system is mounted 

on the top center of the test pile or on a bracket mounted along the line of load 

application.

It is strongly recommended that lateral deflection measurements versus depth also 

be obtained during a lateral load test. This can be accomplished either by installing 

an inclinometer casing on or in the test pile to the depth where it is reasonable 

certain that no lateral movement will occur and recording inclinometer readings 

immediately after application or removal of a load increment held for a duration of 30 

minutes or longer. Alternatively, a Shape Accel Array (SSA) can be embedded in or 

attached to the test pile. Kyfor et al. (1992) noted that lateral load tests in which only 

the lateral deflection of the pile head is measured are seldom justifiable. Additional 

details on load application, and pile head load and movement measurements may 

be found in ASTM D3966 and FHWA-SA-91-042 (Kyfor et al. 1992).

A photograph of a typical lateral load test arrangement is presented in Figure 9-12. 

This figure shows a 14 and a 16 inch O.D. concrete filled pipe pile being pushed 

apart. The jack is located adjacent to the right pile and the load cell and spherical 

bearing plate are located adjacent to the left pile. Figures 9-13 and 9-14 present 

close-ups of these devices. Both piles were also equipped with a string of in-place 

inclinometers for prompt readout of the deflected pile shape with each load 

increment. A photograph of the multiple inclinometer string components is 

presented in Figure 9-15. 

As an alternative means to measure horizontal deflection, a Shape Accel Array 

(SSA) can be embedded in or placed along the pile length (Rollins et al. 2009). 

These triaxial accelerometers are typically installed in 1 inch diameter PVC casing, 

are recoverable and reusable, and can output displacements for both static and 

dynamic loads (Measureand 2015). SSAs are currently produced in 12 inch 

segments and therefore provide a more detailed assessment of deflections as 

compared with typical inclinometers. Figure 9-16 shows a typical SSA.

In cases where it is not feasible to acquire lateral deflection measurements versus 

depth, it may be useful to measure pile head rotation during the lateral load test. 

The pile head rotation can be obtained by measuring the lateral pile deflection at two 

different elevations above the point of load application or by attaching an 

inclinometer probe or tiltmeter to the pile head. The measured pile head rotation 

may be used to assess appropriate p-y modelling.
26

Figure 9-12 Lateral load test on 14 inch and 16 inch O.D. concrete filled pipe piles

(courtesy of WKG2).

Figure 9-13 Jack for lateral load test (courtesy of WKG2

).
27

Figure 9-14 Spherical bearing plate and load cell for lateral load test

(courtesy of WKG2).

Figure 9-15 Multiple inclinometer string components for lateral load test 

(courtesy of WKG2).
28

Figure 9-16 Typical Shape Accel Array (SSA) (courtesy of Measurand).

9.4.2 Lateral Test Loading Methods

Several loading procedures are detailed in ASTM D3966. The Standard Loading 

procedure requires that the total test load be 200% of the proposed lateral design 

load. Variable load increments are applied with the magnitude of load increment 

decreasing with applied load. The load duration is also variable, increasing from 10 

minutes early in the test to 60 minutes at the maximum load. Upon completing the 

maximum test load, the pile is unloaded in four load decrements equal to 50% of the 

maximum load, with 10 minute holds between load decrements. This loading 

schedule is shown in Table 9-2. 
29

Table 9-2 Standard Loading Schedule (after ASTM D3966)

Percent of Design Load Load Duration, minutes

0 - 

25 10

50 10

75 15

100 20

125 20

150 20

170 20

180 20

190 20

200 60

150 10

100 10

50 10

0 30

Readings of load and gross movement are recorded immediately before and after 

each change in load, with additional readings taken at 5 minute intervals between 

load increments. Upon load removal, readings should be taken at 15 and 30 

minutes.

9.4.3 Presentation and Interpretation of Lateral Test Results

The results of lateral load tests should be presented in a report conforming to the 

requirements of ASTM D3966. The interpretation and analysis of lateral load test 

results is much more complicated than those for compression and tension load tests. 

Figure 9-17 presents a typical pile head load-movement curve from a lateral load 

test. A lateral deflection versus depth curve similar to the one shown in Figure 9-18

should also be reported for lateral load tests that include lateral deflection 

measurements versus depth. The measured lateral deflection behavior versus 

depth should also be plotted and compared to design analysis results from p-y 

computer software as indicated in Figure 9-19. Based upon the comparison of 

measured and predicted results, the p-y curves to be used for design (design stage 

tests), or the validity of the p-y curves on which the design was based (construction 

stage tests) can be determined.
30

Refer to FHWA IP 84 11, Handbook on Design of Piles and Drilled Shafts Under 

Lateral Load by Reese (1984) as well as FHWA-SA-91-042, Static Testing of Deep 

Foundation by Kyfor et al. (1992) for additional information on methods of analysis 

and interpretation of lateral load test results.

Figure 9-17 Typical lateral load test, pile head load-deflection curve.
31

Figure 9-18 Plot of lateral load test measured deflected shape versus depth.

Figure 9-19Comparison of measured and COM624P predicted load deflection 

behavior versus depth (after Kyfor et al. 1992).
32

9.5 LOAD TRANSFER EVALUATIONS

The magnitude of shaft and toe resistances or the unit shaft resistance values can 

be determined though instrumented static load tests. Load transfer evaluation along 

the pile shaft may be determined through the use of telltale rods (extensometers) or 

through surface mounted or embedded strain gages. Details on the use of telltales 

or strain gages for load transfer determinations are discussed in the following 

sections. 

9.5.1 Use of Telltales

Telltales are thin steel rods that extend from the pile head to a selected point in the 

pile. The rods are encased within a slightly larger tube. A schematic of telltale rod 

placement within a load test is shown in Figure 9-20. The NYSDOT Static Pile Load 

Test Manual, GCP-18 (2007) provides telltale details for pipe piles, H-piles, and 

timber piles.

LVDT’s or dial gages attached to the top of the telltale rod measure the relative 

movement between the rod attachment locations on the pile and other points such 

as the independent reference beam or the pile head. When using multiple rods, the 

deflection difference between the rod attachment points can provide the load 

transfer between those fixed locations. According to Kyfor et al. (1992), the average 

load in the pile, Qavg, between two measuring points can be determined as follows:

L

RR

Qavg AE − 21 = Eq. 9-4 

Where:

Qavg = average load in pile between two points (kips).

A = pile cross sectional area (in2).

E = elastic modulus of pile material (ksi).

R1 = deflection reading at upper measurement location (inches).

R2 = deflection reading at lower measurement location (inches).

L = length of pile between two measuring points under no load condition.

If the R1 and R2 readings correspond to the pile head and the pile toe respectively, 

then an estimate of the shaft and toe resistances may be computed. For a pile with 

an assumed constant soil resistance distribution (uniform), Fellenius (1990) states 
33

Figure 9-20 Diagram of telltale rods installed on pile 

(modified from Kyfor et al. 1992).

that an estimate of the toe resistance, Rp, can be computed from the applied pile test 

load, Q. The applied pile head load, Q, is chosen as close to the nominal resistance 

or failure load as possible.

 Eq. 9-5 

Where:

Rp = nominal toe resistance (kips).

Qavg = average load in pile between two points (kips).

Q = test load, applied at the pile head (kips).

For a pile with an assumed linearly increasing soil resistance distribution (triangular), 

the estimated toe resistance may be calculated using:

 Eq. 9-6 
34

Where:

Rp = nominal toe resistance (kips).

Qavg = average load in pile between two points (kips).

Q = test load, applied at the pile head (kips). 

The estimated shaft resistance can then be calculated from the applied pile head

load minus the toe resistance.

9.5.2 Use of Strain Gages

When detailed load transfer data is desired, telltale measurements alone are 

insufficient, due to the presence of unaccounted for residual load. Dunnicliff (1988) 

suggests that weldable vibrating wire strain gages be used on steel piles and sister 

bars with vibrating wire strain gages be embedded in concrete piles for detailed load 

transfer evaluations. Other gage types may also be used if load transfer 

measurements under both static and dynamic conditions are desired. A 

geotechnical instrumentation specialist should be used to select the appropriate 

instrumentation for the application, to select instrumentation that can withstand pile 

handling and pile driving, to determine the needed instrumentation redundancy, and 

to determine the appropriate data acquisition system for the program.

A sister bar vibrating wire strain gage is presented in Figure 9-21. The protective 

cover has been removed from the top gage in the box. Sister bar gages are often 

cast into prestressed concrete piles or embedded in concrete filled pipe piles during 

concrete placement. The sister bars are tied to the longitudinal rebar in the casting 

yard for prestressed concrete piles or attached to centralized steel bar or pipe when 

cast into a concrete filled pipe pile.

An arc-weldable vibrating wire strain gage attached to a steel H-pile is shown in 

Figure 9-22. The gage attachment blocks are welded to the pile and then the 

vibrating wire gage positioned between the mounting blocks. 

Piles can also be instrumented with resistance type strain gages. These gages can 

be used for both static load-transfer and dynamic measurements on the pile. A bolton, waterproof, foil resistance strain gage attached to the side of a steel pipe pile is 

shown in Figure 9-23. The gages and instrumentation cables are covered and 

protected by a steel channel as the pile is driven below grade. Resistance type 

strain gages mounted on sister bars are shown in Figure 9-24. These gages are 

being cast into a prestressed concrete pile. The center sister bar also includes an 

accelerometer for dynamic testing purposes.
35

Figure 9-21 Vibrating wire strain gage sister bars for concrete embedment. 

Figure 9-22 Vibrating wire strain gage with welded anchor blocks and 

protective channel. 
36

Figure 9-23 Waterproof electrical resistance strain gage bolted on pipe pile 

with protective channel (courtesy of Besix).

Figure 9-24 Electrical resistance strain gage on sister bars in concrete pile 

casting bed.
37

Moisture is one of the leading causes of resistance strain gage failure. Therefore, all 

lead wires should be carefully checked for any nicks or abrasion and sealed as 

appropriate in these instrumentation situations. A multiple channel data acquisition 

system is also required as part of the instrumentation system.

Strain gages are generally attached at pre-selected points along the pile length to 

determine load-transfer and unit shaft resistances in specific strata or at prescribed 

locations. The closest soil boring to the load test location should be reviewed as 

part of the load-transfer instrumentation program planning so that the desired unit 

shaft resistance values for the selected geomaterials are determined. Additional 

gages can be distributed based upon the Engineer’s experience, the need for 

instrumentation redundancy, site specific conditions, and project requirements. Cost 

effective designs can be finalized based on the determined geomaterial resistances.

The uppermost gage location should be below the pile head in an area where shaft 

resistance does not act on the pile. Dunnicliff (1988) recommends three undamaged 

pile diameters be left above the ground surface to simulate an unconfined 

compression specimen. At a minimum, the pile head strain gages should be at least 

two pile diameters below the pile head to allow for full load development across the 

pile section. At this location, a modulus determination can be made while a 

comparison of internal and external strain readings can indicate composite section 

action for concrete filled pipe piles (Sellers 1995; Komurka 2015). As depicted in 

Figure 9-25, it is also prudent to mitigate bending effects by using multiple external 

gage pairs equidistant from the head and center of the pile, or centralizers on single 

embedded gages. 

Komurka (2015) also recommended another gage location; two pile diameters above 

the pile toe to be used for load transfer evaluation near the pile toe. He reported 

load transfer can be extrapolated from this location over the remainder of the pile to 

estimate the pile toe resistance. Sister bars near the pile toe should not rest on the 

toe when cast into concrete filled pipe piles but should instead be supported from the 

pile head during the concrete curing process.
38

Figure 9-25 Multiple externally mounted strain gages (2 on each web face) 

located in soil resistance free area during static load test (courtesy WKG2).

Following strain gage attachment and pile installation, the axial force in the plane of 

the gage can be determined using Eq. 9-7.

 Eq. 9-7 

Where:

F = axial force in plane of gage (kips).

ε = strain measured in gage.

E = elastic modulus of pile material (ksi).

A = pile cross sectional area (in2).

The modulus of elasticity for steel is well defined. However, for concrete filled steel 

piles, a composite modulus of the concrete and steel section must be calculated 

using Eq. 9-8.

 

Eq. 9-8 
39

Where:

E = elastic modulus of pile material (ksi).

Est = elastic modulus of steel (ksi).

Ast = cross sectional area of steel (in2).

Ec = elastic modulus of concrete (ksi).

Ac = cross sectional area of concrete (in2).

The modulus of elasticity of concrete also decreases with increasing strain, and this 

value should vary over the pile length due to load-transfer. Fellenius (2001), 

recommended the Tangent Modulus Method be used to determine a secant 

modulus, Esm, for load transfer evaluations of concrete piles. The tangent modulus 

versus measured microstrain is plotted for all strain gage levels and all load 

increments. As shown in Figure 9-26, a best fit line through the data determines the 

slope, A, and y-axis intercept, B. Following determination of the tangent modulus 

line from Eq. 9-9, integration is used to calculate stress in Eq. 9-10. 

M t  += BA 

 





∆

∆ = ε

ε

σ Eq. 9-9 

Bεεσ

A + 

 



 = 2

2 Eq. 9-10

Where:

Mt = tangent modulus (ksi).

Δσ = change of stress from one load increment to the next (ksi).

Δε = change of strain from one load increment to the next.

A = slope of tangent modulus line.

ε = strain measured in gage.

B = Y-intercept of the tangent modulus line.

σ = stress (ksi).

Through the use of Hooke’s law, the secant modulus is related to stress and strain in 

Eq. 9-11. By substituting terms from Eq. 9-10, the secant modulus can be 

expressed as shown in Eq. 9-12. 

= Esmεσ Eq. 9-11

sm 5.0 ε += BAE Eq. 9-12
40

Where:

σ = stress (ksi).

Esm = secant modulus (ksi).

ε = strain measured in gage.

A = Slope of tangent modulus line.

B = Y-intercept of the tangent modulus line.

Figure 9-26 Tangent modulus method for determining elastic modulus 

(modified from Fellenius 2001).

Once the modulus for the pile material is determined, Eq. 9-7 can be used to 

determine the force or load at a given depth. A plot of axial load versus elevation 

from a static load test is presented in Figure 9-27. In this figure, seven embedded 

gages were installed to provide a detailed load transfer evaluation for a 14 inch 

diameter, concrete filled, steel pipe pile. The load transferred to the soil can be 

calculated from the difference in force between selected measurement locations. For 

example, when utilizing the final load increment, 16.3 tons of load (e.g. 53.9 tons - 

37.6 tons equals 16.3 tons) is transferred to the soil between the strain gages 

located at elevations EL 549.7 and EL 537.4. The pile in Figure 9-27 is 3.67 feet in 
41

circumference, and because the final two strain gages are located 12.3 feet apart, 

the unit shaft resistance over this interval is 16.3 tons divided by 45.1 ft2 (12.3 feet 

times 3.67 feet) or 0.36 tsf. It should be noted that this computation did not consider 

the effects of any residual load in the pile which is discussed in Section 9.5.3.

Figure 9-27 Plot of axial load versus strain gage elevation for load transfer 

assessment. 

Unit shaft resistance values can be used to evaluate nominal resistances of other 

similar size and type of piles (e.g. 12 inch or 16 inch O.D. in the above case) or to 

refine the pile penetration depth for the required nominal resistance. When 

performed in the design phase, foundation design can be optimized through a more 

complete understanding of load transfer mechanisms in the geomaterials.

9.5.3 Determination of Residual Load

During driving, residual loads can be locked into a pile that does not completely 

rebound after a hammer blow (i.e. return to a condition of zero stress along its entire 

length). This is particularly true for flexible piles, piles with large shaft resistances, 

and piles with large toe quakes. Load transfer evaluations using telltale 

measurements described above assume that no residual loads are locked in the pile 
42

during driving. Therefore, the load distribution calculated from Eq. 9-5 to 9-7 would 

not include residual loads. If measuring points R1 and R2 in Figure 9-20 correspond 

to the pile head and pile toe of a pile that has locked-in residual loads, the calculated 

average pile load would also include the residual loads. This would result in a lower 

toe resistance being calculated than actually exists as depicted in Figure 9-28. 

Figure 9-28 Example of residual load effects on load transfer evaluation.

In a static load test performed on a pile with residual loads, negative shaft resistance 

will be present along the upper portion of the pile. Figure 9-28 illustrates the minimal 

residual load in the upper portion of the pile which must be overcome before the true 

shaft resistance from the applied load can be measured. Moreover, residual loads 

are present along the pile shaft and at the pile toe before the static load test begins. 

This concept is illustrated for mobilized shaft resistance in Figure 9-29, Fellenius 

(2014). If no residual loads acted upon the pile prior to static load testing, Path O-BC would represent the developed shaft resistance. However when residual loads 

are present, Point D is the origin of the residual load, which initially loads the pile 

along Path D-A. Thus, when applying load during the static load test, shaft 

resistance develops along Path A-O-B-C. If residual load is not accounted for, the 
43

loading path will be thought to move through Path A-B-C, and a shaft resistance 

twice as large as the true shaft resistance will be measured, Fellenius (2014).

Figure 9-29 Mobilized shaft resistance hysteresis loop for pile during static 

load test (after Fellenius 2014).

Residual load can also develop at the pile toe. In Figure 9-30, Fellenius (2014), 

presents paths for toe resistance mobilization. Without the presence of residual 

load, toe resistance is developed along Path O-B-C. However when residual load 

exists within the pile, load develops along Path A-D-B-C. Toe resistance therefore 

develops prior to applying load during the static load test. In the above mentioned 

case of unloading, toe resistance is unloaded following the hysteresis loop of Path 

B-D’-A (when starting at Point B), where subsequent loading follows Path A-D-B-C. 

In summation, the residual load can result in underestimating the developed toe 

resistance.
44

Figure 9-30 Mobilized toe resistance for pile during static load test. 

(after Fellenius 2014).

It is apparent from the above discussion that residual load locked into a pile can 

affect the apparent resistance distribution determined from a static load test. While 

the geotechnical nominal resistance is unaffected, the computed shaft and toe 

resistances are altered. Therefore, additional measures should be taken to 

determine the true resistance distribution and unit resistance values. This is 

particularly important when uplift resistance is being assessed from the shaft 

resistance determined in an instrumented compression load test or where unit 

resistance values are being used for design decisions.

Fellenius (2002b) presented a case where residual loads were evaluated from a 

strain gage instrumented static load test in conjunction with subsurface exploration 

results. When using strain gages for load-transfer in a static load test, it is important 

that the” zero reading” be made prior to pile installation (and for sister bars, the 

factory reading) such that no load exists in the pile. Fellenius et al. (2003) noted that 

a lower bound estimate of residual load may be obtained from the difference in strain 

measured just before commencement of the load test and at the true no load 

condition. This calculation of the residual load may however be an underestimate, in 

particular, when thermal strains occur in hydrating concrete.
45

Figure 9-31 presents strain gage instrumented static load test results for a 

prestressed concrete pile driven 48 feet into a loose to medium dense uniform sand. 

Reference should be made to Altaee et al. (1992) and Fellenius (2002b) for further 

details on the static load test or subsurface conditions. The “measured load” at a 

given depth is calculated using Eq. 9-7 and then that load subtracted from the 

applied pile head load to obtain the “measured load” versus depth.

In Figure 9-31 the measured load curve has an inflection point near a depth of 23 

feet, and progresses downwards with increasing slope. This should not be the case 

in a uniform subsurface profile as it would require the unit shaft resistance to be 

smaller in the bottom third of the pile than in the middle portion of the pile. 

Therefore, residual load is likely to have influenced the true load distribution.

Figure 9-31 Instrumented static load test results and soil resistance distribution

(modified from Fellenius 2002b with data from Altaee et al. 1992).
46

After plotting the measured load and calculating unit resistances along the pile, the 

measured shaft resistance divided by 2 (from the Measured Load) versus depth is 

added to the Figure 9-31 graph. The curve for the True Resistance then follows 

down to the inflection level (calculated as the difference in load applied at the pile 

head and the Measured Resistance divided by 2). Below this point, and because 

the uniform soil should exhibit similar shear strengths, the True Resistance curve 

may be extrapolated downward. In effect, it is assumed that the remaining unit shaft 

resistance is equal to that in the length just above this inflection, and does not follow 

the “false” measured values that indicate otherwise. The residual load is plotted as 

the difference of the True Resistance (with the extrapolated portion) and Measured 

Load. Engineering judgement and careful interpretation of the subsurface profile 

and load test results is necessary when determining residual load.

9.6 PRACTICAL ISSUES AND CONSIDERATIONS

Static load tests are the most reliable and most expensive method for determination 

of nominal geotechnical resistance or lateral deformation. Static load tests are 

sometimes avoided because of cost concerns or potential time delays in design or 

construction. While the economic benefits of performing a static load test should be 

carefully considered, cost alone should not be the deciding factor. Load transfer 

information from axial tests or deflected shape versus depth behavior in lateral load 

tests is extremely valuable when performed and evaluated by knowledgeable 

instrumentation specialists who should be involved in static load test program 

development and implementation.

Delays to a project in the design or construction stage usually occur when the 

decision to perform static load tests is added late in the project. During a design 

stage program, delays can be minimized by determining early in the project whether 

a static load test program should be performed. In the construction stage, delays 

can be minimized by clearly specifying the number and locations of static load test to 

be performed as well as the time necessary for the engineer to review the results. In 

addition, the specifications should state that the static load test must be performed

prior to ordering pile lengths or commencing production driving. In this way, the test 

results are available to the design and construction engineer early in the project so 

that the maximum benefits can be obtained. At the same time the contractor is also 

aware of the test requirements and analysis duration and can schedule the project 

accordingly.

Fellenius (1984) reported on a static load test where the initial test was performed 

with the load determined from a load cell and jack pressure gage. In this test, the 
47

load determined from the load cell exceeded the load from the jack pressure gage 

and the accuracy of the test was in question. A second load test was performed 

after replacing the load cell and recalibrating the jack and load cell system. A 

comparison of the load from the jack and load cell in the second test is shown in 

Figure 9-32. The load determined from the jack pressure gage exceeded the load 

determined by the load cell by 10 to 20% in loading and underestimated the load by 

5% in unloading. When load cell and jack pressure readings do not agree, the 

source of the discrepancy should be determined and the load test rerun if needed.

Figure 9-32 Variation between applied pile head load determined from load 

cell and jack pressure gage (after Fellenius 1984).

Load cells and jacks should also be properly sized to reduce measurement errors in 

the applied load. Geokon (2013) noted load cell error typically occurs if the load cell 

and hydraulic jack piston are not equal in diameter. The thickness of the bearing 

plates can further compound loading error, especially if the plate(s) is not sufficiently 

thick to reduce bending effects. In the Geokon study, three jack piston sizes were 

used while retaining the same load cell of size 4 inch I.D., 5-3/4 inch O.D with a 

maximum load of 300 kips. Furthermore, electrical resistance strain gages were 

bonded to the load cell’s middle outside circumference, Sellers (2015). Jack A had 
48

dimensions of 2 inch I.D., 4 inch O.D., Jack B had dimensions of 4 inch I.D., 5-3/4 

inch O.D., and Jack C had dimensions of 6 inch I.D, 8 inch O.D. Figure 9-33

summarizes size effects for the jack and bearing plate to the load cell response. 

Figure 9-33 Load cell response based on jack and bearing plate size 

(after Geokon 2013).

9.7 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS

The advantages of performing static load tests are summarized below. 

1. A static load test allows a more rational design. Confirmation of geomaterial 

resistance through static load testing is considerably more reliable than 

resistance estimates from static analyses and dynamic formulas.

2. An improved knowledge of pile-soil behavior is obtained that may allow a 

reduction in pile lengths or an increase in the factored pile load, either of 

which may result in potential savings in foundation costs.

3. With the improved knowledge of pile-soil behavior, a higher resistance factor 

may be used on the determined nominal resistance. Resistance factors 

between 0.75 and 0.80 are recommended when static load tests are utilized, 

as compared to a resistance factor of 0.40 when using the Modified Gates 

dynamic formula. Hence, a cost savings potential again exists.
49

4. The nominal resistance determined from load testing allows confirmation that 

the factored load may be adequately supported at the planned pile 

penetration depth.

Disadvantages include the load test cost, the time required to setup and perform a 

test, as well as the time required for complete result interpretation (particularly for 

tests with load transfer data). In addition, minimal information is obtained on driving 

stress levels or on the extent of pile damage (if any). Static load test results can 

also be misleading on projects with highly variable soil conditions. Site size and 

variability should therefore be considered when relying upon results for a single test 

pile.

As for limitations, static load test results cannot be used to account for long-term 

settlement, downdrag from consolidating and settling soils, or to adequately 

represent pile group action.
50

REFERENCES

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Highway Administration, Washington, D.C., 82 p.

Abu-Hejleh, N., Abu-Farsakh, M., Suleiman, M.T., Tsai, C. (2015). State Practices 

in Databases for Deep Foundation Load Tests. GSP 256: Proceedings of the 

International Foundations Conference and Equipment Exposition 2015. San 

Antonio, TX, pp. 237-246.

Altaee, A., Fellenius, B.H., and Evgin E. (1992). Axial Load Transfer for Piles in 

Sand. I: Tests on an Instrumented Precast Pile. Canadian Geotechnical 

Journal, Vol. 29, No. 1, pp. 11-20.

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(2014). AASHTO LRFD Bridge Design Specifications, US Customary Units, 

Seventh Edition, with 2015 Interim Revisions. American Association of State 

Highway and Transportation Officials, Washington, D.C., 1960 p.

ASTM D1143-07. (2014). Standard Test Methods for Deep Foundations Under 

Static Axial Compressive Load. Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 15 p.

ASTM D3689-07. (2014). Standard Test Methods for Deep Foundations Under 

Static Axial Tensile Load. Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 13 p.

ASTM D3966-07. (2014). Standard Test Methods Deep Foundation Under Lateral 

Load. Book of ASTM Standards, Vol. 4.08, ASTM International, West 

Conshohocken, PA, 18 p.

Cheney, R.S. and Chassie, R.G. (2000). Soils and Foundations Workshop 

Reference Manual. FHWA HI-00-045, U.S. Department of Transportation, 

National Highway Institute, Federal Highway Administration, Washington, 

D.C., 358 p.
51

Crowther, C.L. (1988). Load Testing of Deep Foundations: the Planning, Design, 

and Conduct of Pile Load Tests. John Wiley & Sons, New York, NY, 233 p.

Davisson, M.T. (1972). High Capacity Piles. Proceedings, Soil Mechanics Lecture 

Series on Innovations in Foundation Construction, American Society of Civil 

Engineers, ASCE, Illinois Section, Chicago, IL, pp. 81-112.

Dunnicliff, J. (1988). Geotechnical Instrumentation for Monitoring Field 

Performance. John Wiley & Sons, New York, NY, pp. 467-479.

Fellenius, B.H. (1984). Ignorance is Bliss – and That is Why We Sleep so Well. 

Geotechnical News Magazine, Vol. 2, No. 4, pp. 14-15.

Fellenius, B.H. (1990). Guidelines for the Interpretation of the Static Loading Test. 

Deep Foundations Institute Short Course Text, First Edition, 44 p.

Fellenius, B.H. (2001). From Strain Measurements to Load in an Instrumented Pile. 

Geotechnical News Magazine, Vol. 19, No. 1, pp. 35-38.

Fellenius, B.H. (2002a). Determining the Resistance Distribution in Piles. Part 1: 

Notes on Shift of No-Load Reading and Residual Load. Geotechnical News 

Magazine, Vol. 20, No. 2, pp. 35-38.

Fellenius, B.H. (2002b). Determining the Resistance Distribution in Piles. Part 2: 

Method for Determining the Residual Load. Geotechnical News Magazine, 

Vol. 20, No. 3, pp. 25-29.

Fellenius, B.H., Harris, D., and Anderson, D.G. (2003). Static Loading Test on a 45 

m Long Pipe Pile in Sandpoint, Idaho. Canadian Geotechnical Journal, 

Vol.41, No. 4, pp. 613-628.

Fellenius, B.H. (2014). Basics of Foundation Design. Electronic Edition. 

http://www.Fellenius.net, 410 p.

Fuller, F.M. (1983). Engineering of Pile Installations. McGraw-Hill, New York, NY, 

286 p.

Geokon (2013) Load Cell Instruction Manual, Appendix C, Load Cell Calibrations – 

Effects of Bearing Plate Warping, Lebanon, NH, pp. 18-20.
52

Komurka, V.E. (2015). Personal communication.

Kyfor, Z.G., Schnore, A.S., Carlo, T.A. and Bailey, P.F. (1992). Static Testing of 

Deep Foundations. Report No. FHWA-SA-91-042, U.S. Department of 

Transportation, Federal Highway Administration, Office of Technology 

Applications, Washington, D.C., 174 p.

New York State Department of Transportation (NYSDOT). (2007). Static Pile Load 

Test Manual, Geotechnical Engineering Bureau, GCP-18, Revision #3, 

http://www.dot.ny.gov, 55 p.

Paikowsky, S.G. (2004), with contributions from Birgisson, B., McVay, M., Nguyen, 

T., Kuo, C., Baecher, G., Ayyub, B., Stenersen, K., O.Malley, K., 

Chernauskas, L., and O’Neill, M., Load and Resistance Factor Design (LRFD) 

for Deep Foundations, NCHRP Report 507. Transportation Research Board, 

Washington, D.C., 76 p.

Reese, L.C. (1984). Handbook on Design of Piles and Drilled Shafts Under Lateral 

Load. Report No. FHWA-IP-84-11, U.S. Department of Transportation, 

Federal Highway Administration, Office of Implementation, McLean, VA, 386 

p.

Rollins, K., Gerber, T., and Cummins, C. (2009). Monitoring Displacement vs. Depth 

in Lateral Pile Load Tests with Shape Accelerometer Arrays, Proceedings of 

the Seventeenth International Conference on Soil Mechanics & Geotechnical 

Engineering, Alexandria, Egypt, Oct. 5-9, 2009. Vol.3, pp. 2016-2019.

Sellers, B. (2015). Personal communication.

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GRLWEAP LRFD Resistance Factor for Oregon DOT. Publication No. ORRD-98-00, Oregon Department of Transportation, Research Unit, Salem, OR, 

82 p.
53

CHAPTER 10

DYNAMIC TESTING AND SIGNAL MATCHING ANALYSIS

Dynamic test methods use measurements of strain and acceleration taken near the 

pile head as a pile is driven or restruck with a pile driving hammer. These dynamic 

measurements can be used to determine the performance of the pile driving system, 

calculate pile installation stresses, assess pile integrity, and evaluate the nominal 

geotechnical resistance.

Dynamic test results should be further evaluated using signal matching techniques 

to determine the relative soil resistance distribution on the pile, as well as dynamic 

soil properties for use in wave equation analyses. This chapter provides a brief 

discussion of dynamic test equipment and analysis methods. 

10.1 BACKGROUND

Work on the development of the dynamic pile testing techniques that have become 

known as the Case Method started with a Master thesis project at Case Institute of 

Technology. This work was done by Eiber (1958) at the suggestion and under the 

direction of Professor H.R. Nara. In this first project, a laboratory study was 

performed in which a rod was driven into dry sand. The Ohio Department of 

Transportation (ODOT) and the Federal Highway Administration subsequently 

funded a project with HPR funds at Case Institute of Technology beginning in 1964. 

This project was directed by Professors R.H. Scanlan and G.G. Goble. The 

research work under the direction of Professor Goble continued to be funded by 

ODOT and FHWA, as well as several other public and private organizations until 

1976.

Four principal directions were explored during the 12 year period that the funded 

research project was active. There was a continuous effort to develop improved 

transducers for the measurement of force and acceleration during pile driving. Field 

equipment for recording and data processing was also continually improved. Model 

piles were driven and tested both statically and dynamically at sites in Ohio. Full 

scale piles driven and statically tested by ODOT, and later other transportation 

agencies, were also tested dynamically to obtain nominal resistance correlations. 

Finally, analysis method improvements were developed, including both field 
54

solutions (Case Method) in the Pile Driving Analyzer system (PDA) and an 

associated signal matching technique (CAse Pile Wave Analysis Program or 

CAPWAP). Additional information on the research project and its results may be 

found in Goble and Rausche (1970), Rausche et al. (1972), and Goble et al. (1975).

ODOT began to apply the results of this research to their construction projects in 

about 1968. Commercial use of the methods began in 1972 when the test 

equipment and analysis methods became practical for use in routine field testing by 

a trained engineer. Further implementation of dynamic testing methods in the 

1980’s resulted from FHWA Demonstration Project 66, in which additional 

correlation data was collected, and method benefits were demonstrated on 

transportation projects throughout the US.

Dynamic testing equipment and signal matching software have continued to be 

improved and enhanced over the 40 years since its original development. Other 

dynamic testing and analysis systems have also developed during that time period, 

primarily in Europe. One of the early European systems was advanced by the 

Netherlands Organization for Applied Scientific Research (TNO). This group 

developed the FPDS equipment and its associated signal matching technique, 

TNOWAVE, Reiding et al. (1988). Additional dynamic testing and analysis systems 

have also emerged over time including the Embedded Data Collector (EDC), 

Hererra et al. (2009), as well as the PDR dynamic testing system and Allwave-DLT 

signal matching software, Middendorp and Verbeek (2010).

10.2 APPLICATIONS FOR DYNAMIC TESTING METHODS

Samtani and Nowatzki (2006) note that dynamic testing costs much less and 

requires less time than static pile load testing. They also note that important 

information can be obtained regarding the behavior of the pile driving system and 

pile-soil response that is not available from a static pile load test. Determination of 

driving stresses and pile integrity with dynamic test methods has resulted in fewer, 

higher nominal resistance piles in foundation designs due to better pile installation 

control. Some of the applications for dynamic testing methods are discussed below.

10.2.1 Nominal Resistance

a. Assessments of nominal geotechnical resistance versus pile penetration 

depth can be obtained by testing from the start to the end of driving. This can 
55

be helpful in profiling the depth to the bearing stratum and thus the required 

production pile lengths.

b. Evaluation of the nominal geotechnical resistance at the time of testing. Soil 

setup or relaxation potential can be assessed by restrike testing several piles 

and comparing nominal resistance at restrike with that at end-of-initial driving.

c. Signal matching analysis provides refined estimates of nominal resistance, 

the soil resistance distribution, as well as insight into soil quake and damping 

parameter selection for wave equation analyses. Signal matching and wave 

equation analysis programs have different pile and soil models. Therefore, 

signal matching determined dynamic soil parameters may require adjustment, 

as described in Section 12.6.9, for input in wave equation analysis programs. 

10.2.2 Hammer and Driving System Performance

a. Calculation of energy transferred to the pile for comparison with the 

manufacturer's rated energy and/or wave equation predictions of hammer and 

drive system performance. Energy transfer can also be used to determine 

the effect of changes in hammer cushion or pile cushion materials on the pile 

penetration resistance or blow count.

b. Determination of drive system performance under different hammer strokes, 

operating pressures, batter angles, or changes in hammer maintenance by 

comparative testing of hammers, or of a single hammer over an extended 

period of use.

c. Identification of hammer performance issues, such as pre-ignition problems 

with diesel hammers or preadmission in air/steam hammers.

d. Determination of whether soil behavior or hammer performance is responsible 

for changes in the observed penetration resistance or blow count.

10.2.3 Driving Stresses and Pile Integrity

a. Calculation of compression and tension driving stresses. In cases with 

driving stress problems, this information can be helpful when evaluating 

adjustments to pile installation procedures. Calculated stresses can also be 

compared to specified driving stress limits.
56

b. Determination of the extent and location of pile structural damage, Rausche 

and Goble (1979). Thus, costly extraction may not be necessary to confirm or 

quantify damage suspected from driving records.

c. Compression and tension stress distribution throughout the pile obtained from 

signal matching. 

10.3 RESISTANCE FACTORS FOR DYNAMIC TESTING 

AASHTO (2014) design specifications provide resistance factors for dynamic pile 

testing with signal matching. When the driving criteria is established by dynamic 

testing with signal matching on two piles per site condition, but no less than 2% of 

the production piles, the AASHTO specified resistance factor is 0.65. When the 

driving criteria is established by dynamically testing with signal matching on 100% of 

the production piles, the AASHTO specified resistance factor is 0.75. The AASHTO 

resistance factor is 0.80 when the driving criteria is established by a successful static 

load test of one pile per site condition in conjunction with dynamic testing with signal 

matching on two piles per site condition but no less than 2% of the production piles. 

When a dynamic test with signal matching is used for determination of the nominal 

resistance in axial tension, the AASHTO specified resistance factor is 0.50. 

AASHTO resistance factors for dynamic testing with signal matching were originally 

developed from data presented in NCHRP Report 507, Load and Resistance Factor 

Design (LRFD) for Deep Foundations, Paikowsky (2004). The database used to 

develop the recommended resistance factors defined the nominal resistance 

determined by a static load test as the load where the static load test load-deflection 

curve exceeded the Davisson offset limit. Similarly, the nominal resistance from 

dynamic testing with signal matching was defined as the nominal resistance during 

restrike determined by the CAPWAP signal matching program. If a different static 

load test interpretation criterion or a different signal matching analysis method is 

used, modification or local calibration of the resistance factor should be considered. 

The Florida Department of Transportation sponsored a research effort by McVay 

and Wasman (2015) to determine the resistance factor for dynamic tests performed 

with the Embedded Data Collector (EDC) system. Resistance factors were 

calculated using First Order Second Moment (FOSM) principles, and using the UF 

and the Tran methods. The research study recommended that the calculated 

resistance values be considered as preliminary due to the limited size of the 

database.
57

10.4 DYNAMIC TESTING

A typical dynamic testing system consists of a minimum of two strain transducers 

and two accelerometers. The reusable gages are externally bolted to diametrically 

opposite sides of the pile at a location two to three diameters below the pile head. 

These gages measure strain and acceleration, and account for non-uniform hammer 

impacts and pile bending. 

Two diametrically opposite mounted strain transducers are required for a valid 

dynamic test to average out and compensate for the influence of non-uniform 

impacts and bending. All driven pile types (prestressed concrete piles; steel pipe, H, 

Monotube, and Tapertube piles; timber piles; and composite piles) can be easily 

tested using external gages with the pile preparation and gage attachment 

procedures varying slightly for each pile type. 

Figure 10-1 illustrates the typical pile preparation procedures required for dynamic 

testing using a reusable external gage system. In Figure 10-1(a), a prestressed 

concrete pile is being prepared for external gage attachment using a hammer drill to 

create holes in the concrete. The holes for the strain transducers are drilled through 

a template to maintain location tolerance. Concrete anchors are then set into the 

drilled holes and the gages are bolted to the concrete anchors. Removal of the 

concrete dust remaining in the drilled holes prior to setting the concrete anchors 

improves anchor bond with the concrete. For steel pipe, Monotube, and Tapertube 

piles, diametrically opposite holes are drilled into the steel as shown in Figure 

10-1(b) and tapped. External gages are then attached using high strength bolts 

inserted in the threaded holes. For steel H-piles, holes are drilled holes through the 

web as shown in Figure 10-1(c) and bolts are used to attach gages on both opposite 

web faces. Wood lag screws are used to attach external gages on timber piles.

Pile preparation and gage attachment typically requires 15 to 20 minutes for each 

pile to be tested. After the gages are attached, the pile driving or restrike process 

continues following usual procedures. For restrike tests, the pile can be drilled and 

gages attached at any convenient location 2 or more diameters below the pile head. 

Drilling near the pile head or reusing the original gage holes at that location is not 

necessary. Most restrike tests are typically on the order of 20 blows or less.
58

Figure 10-1 Pile preparation for dynamic testing. 

A photograph of an externally mounted accelerometer, strain transducer, and Wi-Fi 

transmitter bolted to a steel H-pile is shown in Figure 10-2(a). Signals can be 

transmitted from these gages either via the Wi-Fi transmitter shown, or by a splitter 

cable and main cable that collects and transfers signals from the individual gages. 

System manufacturers also offer a combined strain transducer and accelerometer as 

shown mounted on a pipe pile in Figure 10-2(b). The transmitter or signal collection 

cable is not visible in the photograph. Dynamic test records from either gage 

arrangement are acquired for every hammer blow and transmitted wirelessly or by 

main cable to the data acquisition system.

On concrete piles, dynamic testing can be performed using either reusable external 

gages or embedded gages. An embedded gage set typically consists of one strain 

transducer and one accelerometer cast into the pile at a distance of two to three 

diameters below the pile head. Only one strain transducer and one accelerometer is

required in this situation, provided the embedded gage set is located on the central
59

Figure 10-2 a) Accelerometer, strain gage, and WiFi transmitter (courtesy Pile 

Dynamics) and b) combined strain gage and accelerometer (courtesy Allnamics). 

axis of the pile, such that bending and non-uniform impacts are eliminated. Nonuniform concrete piles, such as those with voided center sections, complicate use of 

embedded gages if the change in cross section occurs near the embedded gage

location. Embedded gages should be located at least one diameter from the cross 

sectional change. An external gage system with the gages placed on the voided 

section below the cross section change is recommended in this situation and for 

concrete cylinder piles where a central axis gage location is not possible.

An embedded gage system consisting of a strain transducer and accelerometer 

being cast into a concrete pile is shown in Figure 10-3. The object on the right side 

sitting atop the casting bed houses the transmitter which will be cast flush mounted 

into the pile surface. The wire on the left side connects the transmitter to a second 

embedded gage set cast into the pile near the pile toe.

Figure 10-4 illustrates another embedded gage arrangement consisting of a sister 

bar mounted strain gage and accelerometer being cast into a concrete pile. Two 

diametrically opposite sister bar mounted strain gages are also shown in the figure. 

The additional diametrically opposite sister bar strain gages are not standard and 

were installed for measurement comparison with the center mounted gage.
60

Figure 10-3 Embedded strain gage and accelerometer unit being cast into concrete 

pile (courtesy Radise International).

Figure 10-4 Embedded resistance strain gage and accelerometer mounted on 

sister bar being cast into concrete pile (courtesy of Pile Dynamics, Inc.).

Embedded gages can also be cast into concrete piles at locations other than near 

the pile head. These additional embedded gages can be monitored during driving 

concurrently with gages located near the pile head for further insight into driving 

stresses, load transfer, or potential pile toe damage.
61

Depending on the dynamic testing system, data can be wirelessly sent from the pile 

to the processing unit using a reusable transmitter as shown in Figure 10-2(a), or 

using a transmitter cast into a concrete pile with the embedded gages.

The data acquisition system conditions and converts the measured strain and 

acceleration signals to force and velocity records versus time. The force is 

computed from the measured strain, ε, times the product of the pile elastic modulus, 

E, and cross sectional area, A, using Equation 10-1.

tF = EAε t)()( Eq. 10-1 

Where:

F(t) = force computed at the gage location at time t (kips).

E = elastic modulus of pile material (ksi).

 A = pile cross sectional area (in2

).

ε(t) = measured strain at time t.

The velocity is obtained by integrating the measured acceleration record over time 

using Equation 10-2.

∫ = )()( dttatV Eq. 10-2 

Where:

V(t) = velocity computed at gage location at time t (ft/s).

a(t) = measured acceleration at gage location at time t (ft/s2).

In most dynamic testing systems, all components for processing, storing, and 

displaying dynamic test signals are combined into either a dedicated field processing 

unit such as the Pile Driving Analyzer (PDA) shown in Figure 10-5 or a laptop 

computer such as the Embedded Data Collector (EDC) shown in Figure 10-6. 

During driving, these systems perform integrations and all other required 

computations to analyze the acquired dynamic records for transferred energy, 

driving stresses, structural integrity, and nominal geotechnical resistance. Numerical 

results for user selected dynamic quantities are also displayed with each blow in real 

time. Basic force and velocity records as well as other results can be viewed on the 

system screen during the test. Processed test records are digitally stored and then 

used for subsequent signal matching analysis performed on-site or in the office as 

well as for graphical and numeric output summaries.
62

Figure 10-5 Pile Driving Analyzer processing unit (courtesy of Pile Dynamics, Inc.).

Figure 10-6 Embedded Data Collector system (courtesy of Radise International). 
63

10.5 BASIC WAVE MECHANICS

This section is intended to summarize basic wave mechanics principles applicable to 

pile driving. Through this general overview, an understanding of dynamic testing 

concepts and how dynamic test results can be qualitatively interpreted can be 

obtained.

When a uniform elastic rod of cross sectional area, A, elastic modulus, E, and wave 

speed, C, is struck by a mass, then a force, F, is generated at the impact surface of 

the rod. This force compresses the adjacent part of the rod. Since the adjacent 

material is compressed, it also experiences an acceleration and attains a particle 

velocity, V. As long as there are no resistance effects on the uniform rod, the force 

in the rod will be equal to the particle velocity times the rod impedance, EA/C. 

Figure 10-7(a) illustrates a uniform rod of length, L, with no resistance effects, that is 

struck at one end by a mass. Force and velocity (particle velocity) waves will be 

created in the rod, as shown in Figure 10-7(b). These waves will then travel down 

the rod at the material wave speed, C. At time L/C, the waves will arrive at the end 

of the rod, as shown in Figures 10-7(c) and 10-7(d). Since there are no resistance 

effects acting on the rod, a free end condition exists. A tensile wave reflection

occurs at a free end which doubles the pile velocity at the free end and the net force 

becomes zero. The wave then travels up the rod with force of the same magnitude 

as the initial input, except in tension, and the velocity of the same magnitude and 

same sign.
64

Figure 10-7 Free end wave mechanics.
65

Consider now that the rod is a pile with no resistance effects, and that force and 

velocity records are obtained from measurements made near the pile head. A 

typical force and velocity record versus time for this "free end" condition is presented

in Figure 10-8. 

Figure 10-8 Force and velocity times (EA/C) records versus time for free end.

The toe response in the records occurs at time 2L/C. This is the time required for 

the waves to travel to the pile toe and back to the measurement location, divided by 

the wave speed. Since there are no resistance effects acting on the pile shaft, the 

force and velocity records are equal until the reflection from the free end condition 

arrives at the measurement location. At time 2L/C, the force wave goes to zero and 

the velocity wave doubles in magnitude. Note the repetitive pattern in the records at 

2L/C intervals generated as the waves continue to travel down and up the pile. This 

illustration is typical of an easy driving situation where the pile "runs" under the 

hammer blow.

Figure 10-9(a) illustrates a uniform rod of length, L, struck by a mass. Again there 

are no resistance effects along the rod length, but the pile end is fixed, i.e., it is 

prevented by some mechanism from moving in such a manner that the particle 

velocity must be zero material wave speed, C. At time L/C, the waves will arrive at 
66

the end of the rod as shown in Figures 10-9(c) and 10-9(d). There the fixed end 

condition will cause a compression wave reflection and therefore the force at the 

fixed end doubles in magnitude and the pile velocity becomes zero. A compression 

wave then travels up the rod.

Consider now that the rod is a pile with a fixed end condition and that force and 

velocity records are again obtained from measurements made near the pile head. 

The force and velocity records versus time for this condition are presented in Figure 

10-10. Since there are no resistance effects acting on the pile shaft, the force and 

velocity records are equal until the reflection from the fixed end condition arrives at 

the measurement location. At time 2L/C, the force wave increases in magnitude and 

the velocity wave goes to zero. This illustration is typical of a hard driving situation 

where the pile is driven to rock. 

As discussed above, the force and velocity records versus time are equal or 

proportional at impact and remain proportional thereafter until affected by soil 

resistance or cross sectional changes. Reflections from either effect will arrive at the 

measurement location at time 2X/C where X is the distance to the soil resistance or 

cross section change. Both soil resistance effects and cross sectional increases will 

cause an increase in the force record and a proportional decrease in the velocity 

record. Conversely, cross sectional reductions, such as those caused by pile 

damage, will cause a decrease in the force record and an increase in the velocity 

record.
67

Figure 10-9 Fixed end wave mechanics.
68

Figure 10-10 Force and velocity times (EA/C) records versus time for fixed end.

The concept of soil resistance effects on force and velocity records can be further 

understood by reviewing the theoretical soil resistance example presented in Figure 

10-11. In this case, the soil resistance on a pile consists only of a small resistance 

located at a depth, A, below the measurement location, and a larger soil resistance 

at depth B. No other soil resistance effects act on the pile, so a free end condition is 

present at the pile toe.

The force and velocity records versus time for this example are presented in the 

lower portion of the figure. The onset of impact occurs when the force and velocity 

records rise together prior to time 0. The time interval from the onset of impact until 

the peak impact velocity is referred to as the rise time. The force and velocity 

records remain proportional or equal until one rise time before time 2A/C. At that 

time, the reflection from the small soil resistance effect begins to arrive at the 

measurement location. This soil resistance reflection causes the small increase in 

the force record and the small decrease in the velocity record at time 2A/C. 

No additional soil resistance effects act on the pile between time 2A/C and time 

2B/C. Therefore, the force and velocity records will remain parallel over this time 

interval with no additional separation. At one rise time prior to time 2B/C, the
69

Figure 10-11 Soil resistance effects on force and velocity records 

(after Hannigan 1990).
70

reflection from the large soil resistance effect will begin to arrive at the measurement 

location. This large soil resistance reflection then causes the large increase in the 

force record and the large decrease in the velocity record at time 2B/C. No 

additional soil resistance effects act on the pile between time 2B/C and time 2L/C. 

Therefore, the force and velocity records exhibit no additional separation over this 

time interval.

At one rise time before time 2L/C, the reflection from the pile toe will arrive at the 

measurement location. Since no resistance is present at the pile toe, a free end 

condition exists, and a tensile wave is reflected. Hence, an increase in the velocity 

record and a decrease in the force record occurs. 

These basic interpretation concepts of force and velocity records versus time can be 

used to qualitatively evaluate the soil resistance effects on a pile. In Figure 

10-12(a), minimal separation occurs between the force and velocity records between 

time 0, or the time of impact, and time 2L/C. In addition, a large increase in the 

velocity record and corresponding decrease in the force record occurs at time 2L/C. 

Hence, this record indicates minimal shaft and minimal toe resistance on the pile.

In Figure 10-12(b), minimal separation again occurs between the force and velocity 

records between time 0 and time 2L/C. However in this example, a large increase in 

the force record and corresponding decrease in the velocity record occurs at time 

2L/C. Therefore, this force and velocity record indicates minimal shaft and a large 

toe resistance on the pile.

In Figure 10-12(c), a large separation between the force and velocity records occurs 

between time 0 and time 2L/C. This force and velocity record indicates a large shaft 

resistance on the pile.
71

Figure 10-12 Typical force and velocity records for various soil resistance conditions.
72

10.6 DYNAMIC TESTING METHODOLOGY

As introduced in Section 10.1, simple field solutions such as the Case Method, and 

more rigorous numerical modeling or signal matching methods such as CAPWAP 

have been developed for analyzing dynamic measurement data. These dynamic 

test methods are sometimes confused with wave equation analysis described in 

Chapter 12, so it is useful to briefly review dynamic test and analysis methods as 

well as their application.

The wave equation is a computer program that is typically performed prior to field 

work. During the design stage, a wave equation drivability analysis is performed to 

check the suitability of potential pile types and sections. During the construction 

phase, the wave equation is used to check the suitability of the contractor’s 

proposed equipment to satisfy nominal resistance and penetration requirements. In 

either application, the program inputs require the engineer to make assumptions on 

the hammer performance and static and dynamic soil response. Following test pile 

installation, a refined wave equation analysis may also be performed. In a refined 

analysis, the engineer uses hammer performance and soil response information 

from dynamic measurements and signal matching results to “calibrate” the wave 

equation to the field conditions. This process is described in Section 12.6.9. The 

wave equation provides a relationship between the nominal geotechnical resistance 

and the pile penetration resistance or blow count. It is therefore often used in 

establishing the driving criteria or in assessing the nominal geotechnical resistance 

of a pile based on its observed pile penetration resistance.

During test pile or production driving, dynamic measurements are made for 

estimates of the nominal geotechnical resistance. During driving, Case Method 

results for nominal resistance are calculated in real time from the measured force 

and velocity records obtained for each hammer blow. The Case Method equations 

for nominal resistance are described in detail in Section 10.6.1. While these simple 

field methods are useful for assessing the nominal resistance, AASHTO (2014) 

requires the nominal resistance be determined through signal matching analysis. 

Additional Case Method equations are used for calculation of driving stresses and 

pile integrity, as well as computation of transferred hammer energy. These 

additional Case Method equations are also described later in this chapter.

Signal matching is a more rigorous numerical analysis procedure that uses the 

measured force and velocity records from one hammer blow. Signal matching 

programs use the dynamic measurement data along with wave equation type pile 

and soil modeling to calculate the nominal geotechnical resistance, the relative soil 
73

resistance distribution, the dynamic soil properties of quake and damping, and 

compression and tension stresses throughout the pile. The nominal resistance 

determined by signal matching is a more accurate assessment of the nominal 

resistance than Case Method results. Signal matching determined soil information 

along with the dynamic test data on driving system performance are often used in 

the development of a refined wave equation analysis. This is the best use of all the 

three methods for driving criteria determination. Signal matching analysis is 

described in greater detail in Section 10.6.6.

10.6.1 Nominal Resistance Determination by Case Method

Research conducted at Case Western Reserve University in Cleveland, Ohio 

resulted in a method which uses electronic measurements taken during pile driving 

to predict nominal resistance. Assuming the pile is linearly elastic and has constant 

cross section, the total static and dynamic resistance on a pile during driving, RTL, 

can be expressed using the following equation, which was derived from a closed 

form solution to the one dimensional wave propagation theory:

 Eq. 10-3 

Where:

RTL= total static and dynamic resistance on pile during driving (kips).

F(t) = force measured at gage location at time t (kips).

V(t) = velocity measured at gage location at time t (ft/s).

t1 = time of first peak input. 

t2 = time of reflection of first peak input from pile toe (tt1 + 2LL/CC).

L = pile length below gage location (feet).

E = elastic modulus of pile material (ksi).

A = pile cross sectional area (in2

).

c = wave speed of pile material (ft/s).

To obtain the static nominal resistance, the dynamic resistance (damping) must be 

subtracted from the above equation. Goble et al. (1975) found that the dynamic 

resistance component could be approximated as a linear function of a damping 

factor times the pile toe velocity, and that the pile toe velocity could be estimated 

from dynamic measurements at the pile head. This led to the standard Case 

Method nominal resistance equation, RSP, expressed as follows:

 Eq. 10-4 
74

Where:

RSP = standard Case Method equation for static nominal resistance (kips).

RTL = total static and dynamic resistance on pile during driving (kips).

J = Case damping factor based on soil type near the pile toe (unit less).

F(t) = force measured at gage location at time t (kips).

V(t) = velocity measured at gage location at time t (ft/s).

t1 = time of first peak input.

E = elastic modulus of pile material (ksi).

A = pile cross sectional area (in2).

c = wave speed of pile material (ft/s).

Typical damping factors versus soil type at the pile toe were determined by finding 

the range in the Case damping factor, J, for a soil type that provided a correlation of 

the RSP static nominal resistance within 20% of the static load test failure load, 

determined using the Davisson (1972) offset limit method. The original range in 

Case damping factor versus soil type from this correlation study, Goble et al. (1975), 

as well as typical ranges in Case damping factor for the RSP equation based on 

subsequent experience, Pile Dynamics, Inc. (2015), are presented in Table 10-1. 

While use of these values may provide good initial estimates of the nominal 

resistance, site specific damping correlations should be developed based upon 

signal matching analysis or static load test results. It should also be noted that Case 

damping is a non-dimensional damping factor and is not the same as the Smith 

damping discussed in Chapter 12 for wave equation analysis.

The RSP or standard Case Method equation is best used to evaluate the nominal 

resistance of low displacement piles, and piles with large shaft resistances. For 

displacement piles driven in soils with large toe quakes and for piles with large toe 

resistances, the maximum toe resistance is often delayed in time. This condition 

can be identified from the force and velocity records. In these instances, the 

standard Case Method equation may indicate a relatively low nominal resistance 

and the maximum Case Method equation, RMX, should be used. The maximum 

Case Method equation searches for the t1 time in the force and velocity records 

which results in the maximum nominal resistance. An example of this technique is 

presented in Figure 10-13. When using the maximum Case Method equation, 

experience has shown that the Case damping factor should be at least 0.4, and on 

the order of 0.2 to 0.4 higher than that used for the standard Case Method equation 

for nominal resistance, RSP. Typical ranges in Case damping factor for the RMX 

equation are also presented in Table 10-1. 
75

Table 10-1 Summary of Case Damping Factors for RSP and RMX Equations

Soil Type at 

Pile Toe

Original Case 

Damping

Correlation Range

Goble et al. (1975)

Recommended

Range in Case 

Damping Constant for 

RSP Equation Pile 

Dynamics (2015)

Recommended

Range in Case 

Damping Constant

for RMX Equation

Pile Dynamics (2015)

Clean Sands 0.05 to 0.20 0.10 to 0.15 0.40 to 0.50 

Silty Sands 0.15 to 0.30 0.15 to 0.25 0.50 to 0.70

Silts 0.20 to 0.45 0.25 to 0.40 0.60 to 0.80 

Silty Clays 0.40 to 0.70 0.40 to 0.70 0.70 to 0.90 

Clay 0.60 to 1.10 0.70 or higher 0.90 or higher

The RMX and RSP Case Method equations are the two most commonly used 

solutions for field evaluation of pile nominal resistance. Additional automatic Case 

Method solutions are available that do not require selection of a Case damping 

factor. These automatic methods, referred to as RAU and RA2, search for the time 

when the pile toe velocity is zero and hence damping is minimal. The RAU method 

may be applicable for piles with minimal shaft resistance, and the RA2 method may 

be applicable to piles with toe resistance plus moderate shaft resistance. These 

automatic methods in their appropriate condition are often helpful supplemental 

indicators of the nominal resistance with the more traditional maximum or standard 

Case Method equations.

While the above Case Method equations are valuable for a quick field assessment of 

the nominal resistance, AASHTO (2014) does not provide a resistance factor for 

these or any other simple direct methods. Signal matching analysis on the collected 

dynamic test data is required for all nominal resistance assessments to be in 

compliance with AASHTO specifications.
76

Figure 10-13 Examples of standard Case Method, RSP, and maximum Case 

Method, RMX, calculations for estimates of nominal resistance. 
77

10.6.2 Soil Resistance Distributions

As noted in Section 10.6, soil resistance effects can be assessed from dynamic test 

records. The relative magnitude of the soil resistance can be evaluated from the 

difference between the force record and velocity times impedance records at a given 

time during the 2L/C time interval following impact. The depth of the soil resistance 

below the gage location can be determined from the reflection time, 2X/C where X is 

the depth and C is the pile wave speed. While relative soil resistance distribution 

effects can be evaluated in this manner, the magnitude of the soil resistance at a 

given depth should be determined from more rigorous signal matching analysis as 

described in Section 10.6.6.

For piles with externally mounted or embedded strain gages that are dynamically 

monitored during driving, the pile forces calculated by measurements at the 

monitored locations can be compared to the calculated forces at those depths in 

signal matching results. However, the presence of residual forces from pile casting 

and/or pile driving greatly complicate any simple closed form solution of the soil 

resistance distribution from externally mounted or embedded instrumentation.

10.6.3 Energy Transfer

The energy transferred to the pile head can be computed from the strain and 

acceleration measurements. As described in Section 10.3, the acceleration signal is 

integrated to obtain velocity, and the strain measurement is converted to force. 

Transferred energy is equal to the work done which can be computed from the 

integral of the force and velocity records over time as given below:

∫ = t

p dttVtFtE

0 )()()( Eq. 10-5 

Where:

Ep(t) = energy computed at the gage location at time t (ft-kips). 

F(t) = force measured at gage location at time t (kips).

V(t) = velocity measured at gage location at time t (ft/s).

This procedure is illustrated in Figure 10-14. The maximum energy transferred to 

the pile head corresponds to the maximum value of Ep(t). The output quantity EMX 

is the maximum value of Ep(t) and can be used to evaluate the performance of the 

hammer and driving system as described in Section 10.7.
78

Figure 10-14 Energy transfer computation.
79

10.6.4 Driving Stresses

The compression stress at the gage location can be calculated using the measured 

strain and pile modulus of elasticity. However, the maximum compression stress in 

the pile may be greater than the compression stress calculated at the gage location, 

such as in the case of a pile driven through soft soils to rock. In these cases, signal 

matching or wave equation analysis may be used to evaluate the maximum 

compression stress elsewhere in the pile. Figure 10-15 illustrates the computation 

process for compression and tension stresses. Force and velocity records for an 18 

inch square prestressed concrete pile with a large toe resistance are presented in 

the top half of the figure. The penetration resistance associated with this record is 

29 blows per inch. The vertical scale between the zero axis and the top of the 

window box is identified as 1500 kips. Point A identifies the maximum compression 

force at the gage location of 795 kips. The maximum compression stress at the 

gage location, CSX, is then this force of 795 kips divided by the pile cross section 

area of 324 in2 or 2.45 ksi.

Computed tension stresses are based upon the superposition of the upward and 

downward traveling waves. The downward traveling wave, wave down identified as 

WD, and the upward traveling wave, wave up identified as WU, are presented in the 

lower half of Figure 10-15. The vertical scale between the zero axis and the top of 

the lower box is once again 1500 kips. The value of wave down, WD, at time (t) is 

computed from the measured force and velocity records according to: 





 



 



 



 +=

c

EA tVtFtWD )()( 2

1 )( Eq. 10-6 

Where:

WD(t) = downward traveling wave, wave down, at time t (kips).

F(t) = force measured at gage location at time t (kips).

V(t) = velocity measured at gage location at time t (ft/s).

E = elastic modulus of pile material (ksi).

A = pile cross sectional area (in2

).

c = wave speed of pile material (ft/s).

The value of wave up, WU, at time (t) is computed from the measured force and 

velocity records according to:





 



 



 



 −= c

EA tVtFtWU )()( 21 )( Eq. 10-7 
80

Where:

WU(t) = upward traveling wave, wave up, at time t (kips).

F(t) = force measured at gage location at time t (kips).

V(t) = velocity measured at gage location at time t (ft/s).

E = elastic modulus of pile material (ksi).

A = pile cross sectional area (in2).

c = wave speed of pile material (ft/s).

Figure 10-15 Example compression and tension stress calculation.

In Figure 10-15, the tension in the pile is computed from the superposition of the 

maximum upward tension force in wave up at time 2L/C +/- 20% identified by point 

B, and the minimum downward compression force in wave down between time 0 

and time 2L/C identified by point C. For the example presented, these values are 

-62 kips for point B and 45 kips for point C. The computed net tension force, CTN, is 

-17 kips. This corresponds to a computed tension stress maxima within the first 
81

2L/C, TSN, of 0.05 ksi. This low tension stress in the upward traveling wave agrees 

with the hard driving conditions depicted by the force and velocity records and 

reported blow count. 

The maximum tension stress can also occur later in the blow. Therefore, the full 

record length is searched for the minimum net tension force occurring from the 

upward travelling tension wave and the minimum value of the downward travelling 

wave in the previous 2L/C interval. In the example given, the minimum tension force 

occurs as a result of the tension in wave up of -154 kips identified by point D and the 

minimum force in the wave down in the preceding 2L/C interval of -105 kips 

identified by point E. The computed tension force of -259 kips corresponds to a 

computed tension stress, TSX, of 0.80 ksi. Hence, high tension stresses can occur 

in hard driving cases due to the downward traveling wave. This occurs when the 

reflected compression wave from a fixed toe condition reaches the free end at the 

pile head and reflects down the pile as a tension wave.

10.6.5 Pile Integrity

The basic concepts of wave mechanics were presented in Section 10.4. 

Convergence between the force and velocity records prior to the rise time before the 

toe response at time 2L/C indicates a reduction in pile impedance, EA/C. For

uniform cross section piles, an impedance reduction is therefore pile damage.

The Beta Method, developed by Rausche and Goble (1979), is used to assess the 

relative severity of any damage along the pile shaft based on the convergence 

between the force and velocity records. If damage is detected, the relative severity 

of the damage is quantified and assigned a BTA value indicating the approximate 

reduction in pile impedance at the damage location. Rausche and Goble (1979) 

proposed the guidelines in Table 10-2 as in indication of the severity of pile damage. 

Piles with BTA values below 80% correspond to damaged or broken piles.

Table 10-2 Pile Damage Guidelines (after Rausche and Goble 1979)

BTA Severity of Damage

1.0 Undamaged

0.8-1.0 Slightly Damaged

0.6-0.8 Damaged

Below 0.6 Broken
82

A method to assess pile toe damage solely for concrete piles driven using the EDC 

system with an embedded strain gage and accelerometer at both the pile head and 

pile toe was proposed by Verbeek and Middendorp (2011). This method uses a 

change in the prestress strain level in the top or bottom strain gage to assess 

damage in that portion of the pile. A change in the prestress level by more than 50 

microstrains over 10 consecutive blows was recommended as identifying pile 

damage. This method will detect damage occurring due to a change in the prestress 

level but not damage due to concrete tension cracks.

One limitation of the original Beta method was its ability to detect damage near the 

pile toe when high toe resistances and/or stress wave reflections occur. Likins and 

Rausche (2014) revisited the original Beta method based on improvements in signal 

processing since 1979. For concrete piles, the wave speed must be determined 

from an early blow. They proposed near toe damage can then be assessed by 

looking for a reflection occurring too early for the correct 2L/C time. The PDA 

performs this computation and automatically identifies an early toe reflection by a 

BTT value that changes from 100% to 1% thereby highlighting the early toe 

reflection for further study and evaluation by the test engineer. The BTA and BTT 

methods can be used for pile damage assessments on all pile types.

10.6.6 Signal Matching 

Once dynamic test data is acquired, it is routinely analyzed with a signal matching 

software program at select events such as the end of initial driving and beginning of 

restrike, or at key pile penetration depths such as the estimated pile toe elevation. 

Signal matching software programs include CAPWAP, iCAP, TNOWAVE, and AllWave DLT. The pile and soil models used in these programs vary. Hence, both the 

analyst and the end user should fully understand the models contained within a 

specific program as well as that programs performance and limitations under a given 

set of conditions.

Signal matching programs allow a more rigorous evaluation of nominal geotechnical 

resistance, the relative resistance distribution, as well as insight into the soil quake 

and soil damping. Signal matching analyses are typically performed on an individual 

hammer blow selected from the end of driving or beginning of restrike. As such, a 

signal matching analysis refines the field dynamic test results at a particular 

penetration depth or time.

The AASHTO resistance factor for a dynamic test with signal matching was based 

on restrike dynamic test data analyzed with the CAPWAP program. Therefore, the 
83

discussion and examples presented in the remainder of this signal matching section 

will use the analysis procedures and output from that program for illustrative 

purposes.

The CAPWAP signal matching program uses wave equation type pile and soil 

models with the measured force and velocity records from the dynamic test 

replacing the hammer model. In this method, depicted in Figure 10-16, the pile is 

modeled by a series of continuous pile segments and static and dynamic soil 

resistances are modeled by elasto-plastic springs and dashpots, respectively. The 

dynamic test data is used to quantify pile force and pile motion, which are two of the 

three unknowns. The remaining unknown is the boundary conditions, which are 

defined by the soil model.

First, reasonable estimates of the soil resistance distribution, soil quakes, and soil 

damping parameters are made. Then, the measured wave down is used to set the 

pile model in motion. The program then computes the equilibrium wave up, which 

can be compared to the measured wave up. Initially, the computed and measured 

wave up will not agree with each other. Adjustments are made to the soil model and 

the calculation process repeated.

The ability to match the measured and computed waves at various times is 

controlled by different factors. Figure 10-17 presents an initial trial in the signal 

matching process where the measured and computed records do not agree. This 

figure identifies the factors that most influence the signal match quality over a 

particular zone. The assumed shaft resistance distribution has the dominant 

influence on match quality beginning with the rise of the record at time tr before 

impact and continuing for a time duration of 2L/C thereafter. This is identified as 

Zone 1 in Figure 10-17. 
84

Figure 10-16 Schematic of CAPWAP signal matching method.
85

Figure 10-17 Factors most influencing signal matching analysis.

In Zone 2, the toe resistance and toe model (toe damping, toe quake, and toe gap) 

most influence the wave match. Zone 2 begins where Zone 1 ends and continues 

for a time duration equal to the rise time, tr plus 3 ms. During Zone 3, which begins 

where Zone 1 ends and continues for a time duration of the rise time tr plus 5 ms, 

the overall nominal resistance controls the match quality. A good wave match in 

Zone 3 is essential for accurate nominal resistance assessments. Zone 4 begins at 

the end of Zone 2 and continues for a duration of 20 ms. The unloading behavior of 

the soil most influences match quality in this zone.

With each analysis, the program evaluates the match quality by summing the 

absolute values of the relative differences between the measured and computed 

waves. The program computes a match quality number for each analysis that is the 

sum of the individual match quality numbers for each of these four zones. An 

illustration of the iteration process is presented in Figure 10-18. 

Throughout the iteration process, adjustments are made to the soil model until no 

further improvement can be obtained between the measured and computed wave 

up. The resulting soil model from signal matching is then considered the best 

estimate of the nominal resistance, including the soil resistance distribution, the soil 

quakes, and the soil damping characteristics.
86

Figure 10-18 Example of signal matching iteration matching process. 
87

The final graphical results from a signal matching analysis are presented in Figure

10-19. Four plots are included in the graphical output. While wave up matching was 

performed for the analysis, the “final match” plot in the upper left hand corners is 

typically presented in terms of the measured and computed force waves versus 

time. The force and velocity record versus time for the analyzed hammer blow are 

presented in the upper right hand plot. In the lower right hand corner, the shaft 

resistance magnitude on each soil segment in kips/ft is plotted above the load 

transfer diagram in kips. Important numerical results are presented immediately to 

the right of these plots including the analyzed pile properties, the final match quality 

number, maximum compression and tension stresses, as well as dynamic soil 

properties. The final plot, in the lower left hand corner, includes simulated static load 

test result. The pile model along with the soil resistance and quake values are used 

to develop these pile top and pile toe load versus displacement plots. 

Figure 10-19 Example of signal matching graphical output.
88

An example of a signal matching final result summary is presented in Figure 10-20. 

For each soil segment, this table lists the depth below grade and the corresponding 

static soil resistance, Ru. Unit shaft resistance values are also provided in the far 

right hand column and can be compared to expected values from static analyses. 

The shaft and toe quake and damping values are summarized in the “Soil Model 

Parameters / Extensions” table beneath the soil resistance output. The match 

quality number, pile penetration resistance, compression and tension stresses, and 

transferred energy are summarized in the bottom section.

The “EXTREMA TABLE”, presented in Figure 10-21, summarizes the stress 

distribution throughout the pile. This table is important because it indicates if higher 

compression stresses are present elsewhere in the pile below the gage location at 

pile segment number 1. In the example provided, the maximum compression 

stresses is 33.9 ksi and it occurs at a distance of 30.2 feet below the gage location. 

Similarly, the maximum tension stress of 2.56 ksi occurs at a location 77.2 feet 

below the gage location.

Figure 10-22 presents the “CASE METHOD” summary table. This output table can 

be used to determine which Case Method nominal resistance equation and damping 

factor correlates best with the nominal resistance from the more rigorous signal 

matching analysis. Hence, this table helps determine which Case Method equation 

and what damping factor should be used for any similar piles that will not be 

analyzed by signal matching. For the results summarized in Figure 10-22, the RMX 

Case Method equation with a damping factor of 1.45 would likely be selected based 

on correlation with the nominal resistance from the signal matching analysis.

The final output table from signal matching analysis is the “PILE PROFILE AND 

PILE MODEL” table presented in Figure 10-23. For the uniform closed-end pipe pile 

in the analysis example, no significant changes in the pile cross section area, elastic 

modulus, unit weight, or perimeter were entered. For non-uniform pile types of 

differing materials, tapered piles, or open end piles with added impedance due to a 

soil plug, a more complex pile profile would be reported.
89

Figure 10-20 Example signal matching output “Summary Results” table. 
90

Figure 10-21 Example signal matching output “Extrema” table.

Figure 10-22 Example signal matching output “Case Method” table.
91

Figure 10-23 Example signal matching output “Pile Profile and Pile Model” table.

10.7 CONSIDERATIONS IN TEST SPECIFICATION AND IMPLEMENTATION

Dynamic testing is specified in many ways depending upon the information desired 

or purpose of the testing. For example, a number of test piles driven at preselected 

locations may be specified. In this application, the test piles are usually driven in 

advance of, or at the start of, production driving so that the information obtained can

be used to establish driving criteria and/or pile order lengths for each substructure 

unit. Alternatively, or in addition to a test pile program, dynamic tests on a specified 

percentage of the production pile quantity per site condition may be performed. 

Production pile testing is usually performed for quality assurance checks on nominal 

resistance, hammer performance, driving stress compliance, pile integrity, and for 

site variability considerations. Lastly, dynamic testing can be used where it was not

specified to troubleshoot problems that arise during construction such as differing 

penetration depths or unusual driving records.

The number of piles that should be dynamically tested on the project depends upon 

the project size, variability of the subsurface conditions, the availability of static load 

test information, and the reasons for performing the dynamic tests. For example, it 

may be desirable to test a higher percentage of piles where there are difficult 

subsurface conditions with an increased risk of pile damage, or where time 

dependent soil strength changes are being relied upon for a significant portion of the 

nominal resistance.

On smaller projects, AASHTO specifications suggest a minimum of two dynamic 

tests per site condition, but no less than 2% of the production piles. On larger 

projects and small projects with anticipated installation difficulties or significant time 
92

dependent nominal resistance issues, a greater number of piles should be tested. 

Dynamically testing one or two piles per substructure location is not unusual in these 

situations. Regardless of the project size, the design or construction engineer 

should be able to adjust the number and locations of dynamically tested piles based 

on design or construction issues that arise. For example, a change to 100% 

dynamic testing and the resulting higher resistance factor may benefit a project 

having pile length overruns if the cost of testing and associated construction time is 

less than the cost of the pile length overrun.

Restrike dynamic tests should be performed whenever pile nominal resistance is 

being evaluated by dynamic test methods. Restrikes are often specified 24 hours 

after initial driving. However, in fine grained soils, longer time periods are generally 

required for the full time dependent nominal resistance changes to occur. Therefore, 

longer restrike times should be specified in these soil conditions whenever possible. 

On small projects, long restrike durations can present significant construction 

sequencing problems. Even so, at least one longer term restrike should be 

performed in these cases. Longer term restrike should be specified 2 to 6 days after 

the initial 24 hour restrike, depending upon the soil type. A warmed up hammer 

(from driving or restriking a non-test pile) should be used for any restrike test.

In soils that exhibit a large increase in resistance from soil setup, it may not be 

possible to activate all the resistance during restrike with the pile hammer used for 

the pile installation. In this situation, a drop hammer system such as the one shown 

in Figure 10-24 can be used to determine the nominal resistances during restrike. 

Typically, a ram weight of approximately 2% of the desired nominal geotechnical 

resistance is required for resistance mobilization in a high strain dynamic test.

When dynamic testing is performed by a consultant, the requirements for signal 

matching analysis should be clearly addressed in the dynamic testing specification 

including the signal matching method. The AASHTO resistance factor of 0.65 is 

based on signal matching on restrike dynamic test data with the CAPWAP analysis 

method. A modified or locally calibrated resistance factor may be appropriate for 

other signal matching methods. This includes the simplified automatic signal 

matching program, iCAP, that shares some but not all of the soil and pile modeling 

and analysis capabilities of the CAPWAP program.

On design-build projects, it is common for the dynamic testing to be performed under 

the design-build team with independent verification tests or test reviews performed 

by the owner. On conventional design-bid-build projects, it is also sometimes 

contractually convenient to specify that the general contractor retain the services of
93

Figure 10-24 APPLE drop weight system. 

the dynamic testing firm. However, this can create potential problems since the 

contractor is then responsible for the owner’s quality assurance program. Some 

agencies have contracted directly with a dynamic testing firm to avoid this potential 

conflict. Other large public owners have acquired dynamic test equipment and 

perform these tests with their in-house staff.

Knowledgeable dynamic testing personnel who properly acquire and interpret 

dynamic test records are important for correct implementation of dynamic test results 

on a project. Therefore, specifications should require that dynamic testing personnel 

attain an appropriate level of expertise on the Pile Driving Contractors Association 

(PDCA) sponsored "Dynamic Measurement and Analysis Proficiency Test" for 

providers of high strain dynamic testing services. The test was designed to reflect 

the knowledge and ability of dynamic test providers which is then indicated in a 

"Certificate of Proficiency." The exam results categorize the six levels of proficiency 

as provisional, basic, intermediate, advanced, master, or expert. This allows 

agencies to specify the level of expertise required for their standard practice or for a 

given project. Additional details on the exam can be found on the PDCA website, 

http://www.piledrivers.org. 
94

10.8 PRESENTATION AND INTERPRETATION OF DYNAMIC TEST RESULTS

The results of dynamic pile tests should be summarized in a formal report that is 

sent to both the construction engineer and foundation designer. The construction 

engineer should understand the information available from the dynamic testing and 

its role in the project construction. As discussed in Chapter 7, numerous factors are 

considered in a pile foundation design. Therefore, the foundation designer should 

review the test results since many considerations; (downdrag, scour, uplift, lateral 

loads, settlement, etc.) may affect the overall design and construction requirements. 

10.8.1 Field Results

Construction personnel are often presented with dynamic testing results with 

minimal guidance on how to use or interpret the information. Therefore, it may be 

helpful to both construction personnel and foundation designers to familiarize 

themselves with a typical screen display and the information available in the field 

during a dynamic test. Figure 10-25 presents a typical dynamic test screen display 

from the Pile Driving Analyzer 8G system.

 

Figure 10-25 Typical dynamic test display from Pile Driving Analyzer. 
95

The main pile information input quantities are displayed near the upper left corner of 

the screen and include the pile length below gages, LE; the pile cross sectional area 

at the gages, AR; the pile elastic modulus, EM; the unit weight of the pile material, 

SP; the pile wave speed, WS; as well as the Case damping factor, JC. 

The graphic portion of the screen is typically divided into two displays as illustrated 

above in Figure 10-25. The upper graphic display presents the force and velocity 

records versus time plotted proportionally. The lower display presents the upward 

travelling wave, WU, and the displacement at the gage location versus time. Both 

displays will change for each hammer blow. The first full height solid vertical line 

represents time t1 in the Case Method calculations and corresponds to the time of 

impact (i.e., the first input peak, as the waves pass the gage location near the pile 

head). The second full height solid vertical line represents time t2 in the Case 

Method calculations and corresponds to the time when the input waves have 

traveled to the pile toe and returned to the gage location at time 2L/C. The partial 

height, dotted, vertical lines preceding the full height lines note the start of the impact 

event or start of the toe reflection, respectively.

An experienced test engineer can visually interpret these signals for data quality, soil 

resistance distribution, and pile integrity. As discussed earlier, soil resistance forces 

cause a relative increase in the force wave and a corresponding relative decrease in 

the velocity wave. Therefore on a pile with a uniform cross section, the separation 

between the force and velocity records between times t1 and t2 indicates the shaft 

resistance. The magnitude of separation is also indicative of the magnitude of the 

total soil resistance above that depth. Toe resistance is indicated by the separation 

between these records beginning at the rise time marker prior to time t2. 

A search for convergence between the force and velocity records is performed 

beginning at the time of the sharp rise in the records prior to time t1 and continuing 

for a time interval of 2L/C thereafter. If convergence between the force and velocity 

records occurs prior to the rise in the velocity record preceding time t2, a cross 

sectional reduction or pile damage is indicated. The degree of convergence 

between the force and velocity records is expressed by the BTA integrity value as a 

percentage of the approximate reduced cross sectional area. As discussed in 

Section 10.6.5, an early reflection prior to 2L/C is also evaluated and, if noted, a BTT 

value of 1% will highlight the need for record review for toe damage by the engineer.

The results of Case Method numerical computations are identified by three letter 

codes displayed on the left hand side of the screen. A summary of the most 
96

commonly computed quantities and their corresponding three letter code is 

presented in Table 10-3. 

Table 10-3 Description of Typical Dynamic Test Output Codes

PDA 

Output Code Output Quantity Description

CSX Maximum compression stress at the gage location.

CSI Maximum compression stress from an individual strain transducer. 

CSB Maximum computed compression stress at pile toe.

TSX Maximum computed tension stress. 

BTA Pile integrity factor. 

BTT Integrity indicator for early toe reflection. 

LTD Length to pile damage.

EMX Maximum energy transferred to the gage location. 

ETR Energy transfer ratio (EMX / E rated).

STK Computed hammer stroke. 

BPM Hammer operating rate.

RMX Maximum Case Method (requires damping factor, J).

RSP Standard Case Method (requires damping factor, J).

RSU Case Method with unloading correction (requires damping factor, J). 

RAU Automatic Case Method - toe bearing. No shaft resistance. 

RA2 Automatic Case Method - Moderate shaft resistance. 

RUC* iCAP nominal resistance. 

SFC* iCAP shaft resistance. 

EBC* iCAP toe resistance. 

CSC* iCAP computed maximum compression stress. 

CBC* iCAP computed compression stress at pile toe. 

TSC* iCAP computed tension stress. 

* - requires additional iCAP automated signal matching software.

In the example given in Figure 10-25, the first four output quantities Q1, Q2, Q3, and 

Q4 provide information on the maximum force and associated driving stresses. The 

maximum average force at the pile head, FMX, is 696 kips. This corresponds to an 

average compression stress at the gage location, CSX, of 44.9 ksi. The maximum 

compression stress from an individual strain transducer, CSI, is 45.3 ksi, and the 
97

maximum computed compression stress at the pile toe, CSB, is 49.5 ksi. Hence, 

driving stress levels are quite high and are above AASHTO recommended limits.

As noted earlier, the compression stress levels exceed the driving stress limits as 

well as the guaranteed minimum yield strength for H-piles made from A572, Grade 

50 steel. Hence, the potential for pile toe damage is high. The pile integrity, BTA, is 

calculated as 100%, indicating that no detected damage. The integrity near the pile 

toe is also assessed and a BTT of 100% is calculated despite the very high driving 

stresses resulting from the abrupt refusal driving on hard rock.

An assessment of the performance of the single acting diesel hammer is made from 

output quantities Q5 through Q9. The average energy transferred to the gage 

location, EMX, is 21.2 ft-kips. As indicated by Q6, this corresponds to an energy 

transfer ratio, ETR, of 49.9% (EMX / manufacturer’s rated hammer energy of 42.4 ftkips). The hammer stroke, STK is 9.42 feet and the hammer operating rate, BPM, is 

38.6 blows per minute.

Output quantities Q11 and Q12 display the nominal resistance as evaluated by the 

maximum Case Method, RMX, with a damping factor of 0.80 and 0.90. These 

nominal resistance computations are 748 and 737 kips, respectively and are 

reported as RX8 and RX9. 

Output quantities Q14 to Q16 present the results of automated signal matching 

analysis using the iCAP method. The iCAP results indicate a nominal resistance, 

RUC, of 568 kips with 28 kips of shaft resistance, SFC, and 539 kips of toe

resistance, EBC. Complete automated signal matching results are presented in the 

upper right corner of the graphical screen and include the match quality, MQ, of 3.90 

along with the maximum computed compression stress, CSC, of 52.3 ksi which 

occurs at the pile toe. It should be emphasized that iCAP automated signal 

matching is only applicable to uniform piles. It cannot accurately analyze nonuniform piles, piles with splice gaps, damaged piles, concrete piles with minor 

cracking, or piles with uncertain properties. iCAP should also not be used for large

diameter open-end pipe piles (due to internal plug movements) or on piles in 

unusual soil conditions.

Construction personnel should review the dynamic test results and check that the 

calculated driving stresses, CSX and TSX, are maintained within specification limits. 

Drive system performance indicated by the transferred energy, EMX, should be 

within a reasonable range of that predicted by wave equation analysis or recorded 

on previous tests at the site. If significant variations in energy are noted, the 
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reasons for the discrepancy should be evaluated. The recorded hammer speed 

should be compared to the manufacturer's specifications. Nominal resistance 

estimates should be compared with the required nominal resistance. In soils with 

time dependent soil strength changes, this comparison of nominal resistance should 

be based on restrike tests and not end of initial driving results.

10.8.2 Evaluation of Hammer and Drive System Performance

The performance of a hammer and driving system can be evaluated from a driving 

system's energy transfer ratio, which is defined as the energy transferred to the pile 

head divided by the manufacturer's rated hammer energy. The energy transfer ratio 

should not be misinterpreted as the hammer efficiency as the energy transfer ratio 

includes all energy losses in the driving system. Numerous factors affect the 

transferred energy and hence the energy transfer ratio. These include the hammer 

stroke, fuel setting, helmet weight, hammer and pile cushions, pile impedance, pile 

length, soil resistance, dynamic soil properties, as well as the hammer efficiency.

Figure 10-26 presents energy transfer ratios for selected hammer and pile type 

combinations expressed as a percentile. In this graph, the average transfer 

efficiency for a given hammer-pile combination can be found by noting where that 

graph intersects the 50th percentile. Depending upon the hammer-pile combination, 

average transferred energies as a percentage of the rated energy range from about 

26% for a diesel hammer on a concrete pile to 69% for a hydraulic hammer on a 

steel pile.

Histograms of the energy transfer ratios for all diesel, single acting (SA) air/steam, 

and single acting (SA) hydraulic hammer on steel or concrete and timber piles are 

presented in Figures 10-27, 10-28, and 10-29, respectively. The histograms may be 

useful in assessing drive system performance as they provided the distribution and 

standard deviation of drive system performance for a given hammer-pile 

combination at the end of drive condition.
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Figure 10-26 Energy transfer ratios for select hammer and pile combinations.
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Figure 10-27 Histograms of energy transfer ratio for diesel hammers on 

(a) steel piles and (b) concrete/timber piles.
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Figure 10-28 Histograms of energy transfer ratio for single acting air/steam 

hammers on (a) steel piles and (b) concrete/timber piles.
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Figure 10-29 Histograms of energy transfer ratio for single acting hydraulic hammers 

on (a) steel piles and (b) concrete/timber piles.
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10.8.3 Test Record Illustrating Problematic Hammer Performance

Records for a single acting diesel hammer exhibiting diesel fuel pre-ignition are 

presented in Figure 10-30. On this project, the diesel hammer had been operated 

relatively continuously all day with minimal down time. Note the magnitude of the 

force record during the pre-compression phase prior to impact, and compare that to 

a more typical diesel hammer record in Figure 10-31. In the presented pre-ignition 

data for Blow 986, the hammer stroke is 7.49 feet, the impact force is 383 kips, the 

transferred energy is 20.6 ft-kips, and the energy transfer ratio is 29.4% which is 

10% less than the mean value for a diesel hammer on a steel pile. In Figure 10-26, 

an energy transfer ratio of 29.4% for a single acting diesel hammer on a steel pile 

falls near the 18th percentile which is also clearly indicative of a problem. For 

comparison, Blow 2 of the same drive sequence had a similar hammer stroke of 

7.53 feet when the diesel fuel was not pre-igniting. However, in Blow 2, the hammer 

had an impact force of 481 kips, a transferred energy of 34.2 ft-kips, and an energy 

transfer ratio of 48.8%.

Figure 10-30 Example dynamic test records on pre-igniting diesel hammer.

Note the magnitude of the force record in Figure 10-30 during the pre-compression 

phase prior to impact. The pre-compression force on the pile is almost 50% of the 
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impact force. This reduces the effectiveness of the hammer blow on the pile 

because it takes more energy to compress the combustion gases in the combustion 

chamber prior to the ram impacting the impact block. The reduced transferred 

energy to the pile results in a greater blow count occurring at a shallower pile 

penetration depth. Hence, if pre-ignition was not detected, pile driving would be 

terminated prematurely at a nominal resistance less than the required resistance.

10.8.4 Test Record Illustrating Pile Damage 

Force and velocity records are presented in the upper graph of Figure 10-31 for a 

HP 14x117 H-pile driven with a single acting diesel hammer. The lower graph 

presents the wave up and displacement records. The pile information section of the 

screen in the left hand corner indicates the H-pile has a length of 88.3 feet below the 

gages. A visual interpretation of the force and velocity record suggests the pile has 

developed moderate shaft resistance over the lower portion of the pile with a 

significant amount of the nominal resistance due to toe resistance. Note that a full 

height, dash and dotted vertical line has also appeared between the two solid 

vertical lines corresponding to the pile head, t1, and pile toe, t2. Convergence 

between the force and velocity records before time 2L/C, as noted by the dash and 

dotted line, indicates a pile impedance reduction or damage. 

The BTA warning box near the top of the screen has also turned black and indicates 

a calculated BTA value of 82%. For the example shown, damage was occurring at a 

depth of 83.2 feet below gages due to the H-pile buckling and bending at this 

location. The integrity near the pile toe has been assessed and a BTT of 1% is 

reported highlighting the occurrence of near toe damage. The pile was extracted to 

confirm the indicated damage 5 feet above the pile toe. Photographs of the 

extracted pile are presented in Figure 10-32. 

10.8.5 Test Records Illustrating Soil Setup 

Test records for a 14 inch O.D. x 0.50 inch wall closed end pipe pile driven with a 

single acting diesel hammer are presented in Figures 10-33 and 10-34. Dynamic 

testing on this test pile was performed during initial driving and again during restrike 

3 days after initial driving. The required nominal resistance was 765 kips based on 

resistance verification using dynamic testing with signal matching. At this test pile 

location, a minimum penetration depth of 90 feet was also specified to satisfy lateral 

loading requirements. Soil conditions consist of predominantly stiff to very stiff silty 

clays with interbedded medium dense to dense silty sand and sandy silt layers.
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Figure 10-31 Example dynamic test records indicating pile damage.

Figure 10-32 Photographs of extracted pile showing damage. 
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Figure 10-33 Example test record illustrating soil setup - end of initial driving data.

Figure 10-34 Example test record illustrating soil setup - beginning of restrike data.
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Figure 10-33 presents a representative dynamic test record obtained near final

driving. The pile penetration resistance or blow count at the end of initial driving was 

9 blows per foot at an average hammer stroke of 5.8 feet. For the presented 

hammer blow, the impact force was 467 kips, the transferred energy was 42.5 ftkips, and the energy transfer ratio was 34.8%. The nominal resistance based on the 

Case Method solution was 154 kips which was well below the 765 kips required. 

This was substantiated by signal matching analysis which indicated a nominal 

resistance of 129 kips, with 75 kips of the resistance carried by shaft resistance and 

54 kips through toe resistance.

Figure 10-34 presents a representative dynamic test record obtained near the 

beginning of restrike 3 days later. The blow count at the beginning of restrike was 8 

blows per inch at an average hammer stroke of 10.1 feet. Hence, both the restrike 

blow count and hammer stroke were significantly higher. For the blow presented, 

the impact force was 866 kips, the transferred energy was 81.0 ft-kips, and the 

energy transfer ratio is 66.5%. The nominal resistance based on the Case Method 

solution was 682 kips, or roughly 4.4 times the nominal resistance at the end of 

driving. A nominal resistance of 737 kips with 472 kips of shaft resistance and 265 

kips of toe resistance was determined by signal matching. Long term restrike tests 

showed that, with additional time, the 765 kip nominal resistance would be obtained.

The increased nominal resistance during restrike is visually apparent when 

comparing the end of initial driving and beginning of restrike test records. The 

separation between the force and velocity records between time 0 and time 2L/C is 

substantially greater in the restrike test data compared to the end of initial driving 

data. This is also apparent in the restrike wave up record which has a much greater 

magnitude as well as significantly steeper slope compared to the end of driving wave 

up record.

The magnitude of the separation between the force and velocity records in the 

restrike data substantially increases beginning near the midpoint between time 0 and 

time 2L/C. This is also apparent in the wave up graph which has a significant 

change in slope occurring at the same time. These records indicate that the much 

larger shaft resistance present during restrike developed primarily over the lower half 

of the pile.

A dramatic change in the test records also occurs at, and immediately, after 2L/C. 

This indicates a change in the toe resistance and dynamic toe response. Signal 

matching analysis indicated an increase in the toe resistance, an increase in both 

the shaft and toe damping, and a reduction in the soil quake at the pile toe.
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10.8.6 Test Records Illustrating Relaxation

Test records for HP 12 x 74 H-pile was driven into very dense, clayey silt with a 

single acting diesel hammer are presented in Figures 10-35 and10-36. SPT N 

values in the clayey silt deposit ranged from 37 to 100 blows per foot with a SPT N 

value of 87 blows per foot closest to the pile toe elevation. Dynamic testing was 

performed on the pile during initial driving and 3 days later during restrike. A

nominal resistance of 408 kips was required based on resistance verification using 

dynamic testing with signal matching.

Figure 10-35 presents a representative dynamic test record obtained near final 

driving. The pile penetration resistance or blow count at the end of initial driving was 

26 blows per foot at an average hammer stroke of 9.1 feet. For the presented 

hammer blow, the impact force was 738 kips, the transferred energy was 38.6 ftkips, and the energy transfer ratio was 55.0%. The nominal resistance based on the 

Case Method solution was 456 kips, exceeding the 408 kips required. This was 

substantiated by signal matching analysis which indicated a nominal resistance of 

471 kips, with 216 kips of the resistance carried by shaft resistance and 255 kips 

through toe resistance. 

Figure 10-36 presents a representative dynamic test record obtained near the 

beginning of restrike. The blow count at the beginning of restrike was 3 blows per 

inch at an average hammer stroke of 8.3 feet. Hence, the restrike blow count was 

slightly higher, but the hammer stroke was reduced. For the blow presented, the 

impact force was 646 kips, the transferred energy was 27.0 ft-kips, and the energy 

transfer ratio was 38.5%. The nominal resistance based on the Case Method 

solution was 351 kips, or roughly 100 kips less than at the end of driving. Signal 

matching analysis indicated a nominal resistance at the beginning of restrike of 330 

kips with 235 kips of shaft resistance and 95 kips of toe resistance. 

The reduced nominal resistance during restrike is evident when comparing the end 

of initial driving and beginning of restrike records. The force and velocity records in 

the top portion of both figures exhibit less separation between force at velocity 

records at, and immediately after, the 2L/C marker. This is also apparent in the 

wave up graphs in the lower portion of both figures. The restrike record has a lower 

magnitude resistance occurring after the 2L/C marker as well as having a flatter 

slope.
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Figure 10-35 Example test record illustrating relaxation - end of initial driving data.

Figure 10-36 Example test record illustrating relaxation - beginning of restrike data.
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10.8.7 Reporting of Dynamic Test Results

Additional insight into the pile and soil behavior during driving can be obtained by 

comparing the dynamic test numerical results versus pile penetration depth and 

corresponding driving resistance. Dynamic testing systems typically assign a 

sequential blow number to each hammer blow. By comparing the pile driving log 

with these blow numbers, numerical and graphical summaries of the dynamic testing 

results versus pile penetration depth and pile penetration resistance can be 

prepared.

An example of a numerical summary of dynamic testing results versus depth for a 24 

inch square prestressed concrete pile driven with an APE D62-22 diesel hammer is 

presented in Figure 10-37. The test pile has a required nominal resistance of 1000 

kips. Compression and tension stresses are limited to 3.80 ksi and 1.50 ksi, 

respectively. The accompanying graphical results are presented in Figure 10-38. 

Specification should require that the dynamic test data for each pile tested be 

processed versus pile penetration depth with the corresponding blow count in a 

similar manner. These numerical and graphical results can easily be compared to 

project requirements by construction personnel.

The effects of fuel setting adjustments as well as pile cushion changes are readily 

apparent in these graphical and numerical results. Near 43 feet, the diesel hammer 

fuel setting was increased from fuel setting 1 to 2. Dynamic test results at this depth 

illustrate an increase in both the compression and tension driving stresses, an 

increase in the hammer stroke, and an increase in transferred energy. A decrease 

in the pile penetration resistance also occurs as a result of the fuel setting 

adjustment even though a small increase in the nominal resistance occurs. Near 68 

feet, the fuel setting is once again increased, this time from fuel setting 2 to 3. At 

that depth and time, the pile cushion is also replaced. When driving is resumed, 

both compression and tension driving stresses increase, the hammer stroke 

increases, and the transferred energy increases. The pile penetration resistance 

decreases as a result of the fuel setting adjustment even though a small increase in 

the nominal resistance once again occurs. Driving is temporarily stopped at the 

estimated pile penetration depth of 76 feet and a one hour restrike performed. 

Unfortunately, the nominal resistance slightly decreases during restrike requiring the 

test pile to subsequently be driven deeper for the required nominal resistance. The 

compression and tension driving stresses were maintained within specification limits 

throughout the test pile installation process.
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Figure 10-37 Typical tabular presentation of dynamic test results.
112

Figure 10-38 Typical graphical presentation of dynamic test results versus depth.

10.9 CASE HISTORY

The following case history illustrates how dynamic pile testing and analysis was 

used on a small single span bridge constructed in a remote area. The subsurface 

exploration for the project found a 98 foot thick deposit of moderately clean, medium 

dense to dense sands with SPT N values ranging from 10 to 50. Based upon these 

conditions, the foundation report recommended 12.75 inch O.D. closed end pipe 

piles be used for the bridge abutment foundations. The pipe piles had an estimated 

length of 39 feet for a nominal resistance of 326 kips. The foundation report 

recommended wave equation analysis be used for construction control. Dynamic 

testing of one test pile at each abutment was also specified with the test pile 

information to be used by the engineer to provide the contractor pile order lengths.

The Case Method was used to evaluate pile nominal resistance versus penetration 

depth during the test pile driving. More rigorous signal matching analyses were also 

performed on the dynamic test data to check the Case Method results at selected 
113

pile penetration depths. During initial driving at Abutment 1, the 12.75 inch pipe pile 

drove beyond the estimated pile penetration depth without developing the required 

nominal resistance. The pile was driven to a depth of 75 feet and had an end of 

drive nominal resistance of 235 kips. A restrike dynamic test performed one day 

after initial driving indicated the nominal resistance increased slightly to 245 kips.

While the test pile information from Abutment 1 was being evaluated, three 

additional test piles were driven at Abutment 2. First, dynamic testing of a 16 inch 

O.D. closed end pipe pile was performed to determine if a larger diameter pipe pile 

could develop the required nominal resistance and, if so what pile penetration depth 

was necessary. The 16 inch pile was driven to a depth of 89 feet and had an end of 

drive nominal resistance of 222 kips. A one day restrike test on this pile indicated a 

nominal resistance of 280 kips. The 16 inch pile was driven deeper following the 

restrike test to a final penetration depth of 112 feet. With the additional driving, the 

nominal resistance at the end of redrive decreased to 240 kips.

Approximately two weeks later, a 12.75 O.D. closed end pipe pile and a 14 inch 

diameter Monotube pile with a 25 foot tapered lower section were driven at 

Abutment 2. The 12.75 inch pipe pile was driven to a penetration depth of 95 feet 

with an end of drive nominal resistance of 105 kips. The Monotube pile was driven 

to a depth of 43 feet and had an end of drive nominal resistance of 190 kips. One 

day restrike tests on both piles indicated a slight increase in nominal resistance to 

180 kips and 205 kips, respectively. During this same site visit, a 16 day restrike 

test was performed on the 16 inch pipe pile. The long term restrike nominal 

resistance for the 16 inch pipe pile was 400 kips.

The dynamic testing results from both abutments indicated that the required nominal 

resistance could not be obtained at or near the estimated pile penetration depth with 

the 12.75 inch pipe piles. However, two foundation solutions were indicated by the 

dynamic testing results. If a reduced nominal resistance were chosen, the test

results indicated a Monotube pile driven to a significantly shorter penetration depth 

could develop about the same nominal resistance as could be developed by the 

12.75 inch pipe piles. Alternatively, if the original nominal resistance was desired, 

16 inch pipe piles could be driven on the order of 92 feet below grade.

Although not originally planned, two static load tests were performed to confirm the 

nominal resistance that could be achieved at the site. The 12.75 inch pipe and the 

14 inch Monotube piles at Abutment 2 were selected for testing. The static load test 

results indicated the 12.75 inch pipe pile with a pile penetration depth of 95 feet had 

a nominal resistance of 230 kips and the Monotube pile with a pile penetration 
114

depth of 43 feet had a nominal resistance of 220 kips. The dynamic test nominal 

resistances determined during restrike were in good agreement with these static 

load tests results particularly when the additional time between the dynamic restrike 

tests and static load tests is considered.

Based on the required pile lengths and the nominal resistances determined from the 

dynamic and static load testing, a cost evaluation of the foundation alternatives was 

performed. The cost analysis indicated that the Monotube piles would be the most 

economical pile foundation type. This case study illustrates how the routine 

application of dynamic testing on a small project helped facilitate the solution to an

unexpected foundation problem.

10.10 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS

An advantage of dynamic testing over other methods of nominal resistance 

verification is the additional information gained on the pile installation process. In 

addition to providing estimates of nominal resistance during driving and during 

restrike, dynamic test data can be used to check hammer and drive system 

performance, to monitor driving stresses, and to assess pile structural integrity.

Many piles can also be dynamically tested during initial driving or during restrike in 

one day. This makes dynamic testing an economical and quick testing method. 

Results are generally available immediately after each hammer blow.

On large projects, dynamic testing can be used to supplement static pile load tests 

or reduce the overall number of static tests to be performed. Since dynamic tests 

are more economical than static tests, additional coverage can also be obtained 

across a project at reduced costs. On small projects where static load tests may be 

difficult to justify economically, dynamic tests offer a viable construction control 

method.

Dynamic tests can provide information on pile nominal resistance versus depth, 

nominal resistance variations between locations, and nominal resistance variations 

with time after installation through restrike tests. This information can be helpful in 

augmenting the foundation design, when available from design stage test pile 

programs, or in optimizing pile lengths when used early in construction test 

programs.
115

When used as a construction monitoring and quality control tool, dynamic testing 

can assist in early detection of pile installation problems such as poor hammer 

performance or high driving stresses. Test results can then facilitate the evaluation 

and solution of these installation problems.

On projects where dynamic testing was not specified and unexpected or erratic 

driving behavior or pile damage problems develop, dynamic testing offers a quick 

and economical method of troubleshooting.

Results from dynamic testing and signal matching analysis can be used to develop 

pile driving criterion. A procedure describing the use of dynamic test results to refine 

wave equation input parameters and wave equation analysis results is described in 

Section 12.6.9.

A disadvantage with dynamic testing for determining the nominal resistance can be

the pile driving system. The pile hammer must be capable of mobilizing all the soil 

resistance acting on the pile. Shaft resistance can generally be mobilized at a 

fraction of the movement required to mobilize the toe resistance. However, when 

pile penetration resistances approach 10 blows per inch, the soil resistance may not 

be fully mobilized at and near the pile toe. In these circumstances, dynamic test 

capacities tend to produce lower bound estimates of the nominal resistance. If 

available, a larger pile hammer or higher hammer stroke can be used to increase the 

net pile penetration per blow and thereby mobilize more resistance, if present.

Dynamic testing estimates of nominal resistance also indicate the nominal 

resistance at the time of testing. Since increases and decreases in the pile nominal 

resistance with time typically occur due to soil setup/relaxation, restrike tests after an 

appropriate waiting period are usually required for a better indication of long term 

pile nominal resistance. This may require an additional move of the pile driving rig 

for restrike testing.

A limitation of dynamic testing can be the geotechnical failure mechanism. Large 

diameter open ended pipe piles or H-piles which do not bear on rock may behave 

differently under dynamic and static loading conditions. This is particularly true if a 

soil plug does not form during driving. In these cases, limited toe resistance 

develops during the dynamic test. However, under slower static loading conditions, 

these open section piles may develop a soil plug and therefore a higher pile nominal 

resistance under static loading conditions. Interpretation of test results by 

experienced personnel is important in these situations.
116

10.11 PRACTICAL ISSUES AND CONSIDERATIONS

Several practical issues and considerations should be clearly understood by the 

parties responsible for analyzing and reviewing dynamic test results. Some of the 

more common issues encountered include:

• Understanding that specifying “a dynamic test” does not implicitly require the 

dynamic test personnel to furnish driving criteria.

Like any other engineering service, dynamic test results should be analyzed 

and reviewed. While it is often desirable to continue driving production piles 

immediately after the dynamic test is complete, time should be allocated for 

analysis and reporting of test results. If the driving criteria is to be determined 

from the dynamic test results, that should be clearly identified in the project 

specifications as well as when the driving criteria is to be furnished to the 

owner or contractor.

• Understanding the limitations of dynamic tests in easy driving and hard driving 

situations.

At pile penetration resistances less than 24 blows per foot and above 120 

blows per foot, dynamic test and analysis methods can overpredict and 

underpredict the nominal resistance, respectively. At low blow counts (high 

set per blow), it is difficult for dynamic methods to easily separate the static 

and dynamic soil resistance effects resulting in a tendency to overpredict the 

static resistance. Use of a reduced hammer stroke or lower fuel setting can 

help improve the accuracy of dynamic methods in low blow count situations. 

At very high blow counts (low set per blow), dynamic test methods tend to 

produce lower bound nominal resistance estimates as not all of the resistance 

(particularly at and near the toe) is fully activated. In these high blow count 

situations, use of a larger hammer stroke, higher fuel setting, pile hammer 

with a greater rated energy, or variable stroke drop hammer can help improve 

dynamic method accuracy.

• Understanding that a dynamic test provides the mobilized nominal resistance at 

the time of testing.

The nominal resistance determined in a dynamic test is the mobilized 

resistance at that particular time. The nominal resistance of a pile typically 

changes over time. It may increase (soil setup) or decrease (relaxation). 
117

Dynamic tests must be performed at both the end of initial driving and during 

restrike at a later time in order to quantify time dependent changes in nominal 

resistance. 

• Knowing how to perform and evaluate restrike dynamic test results.

Ideally, the hammer stroke or fuel setting is selected such that the penetration 

resistance at the beginning of restrike falls between 3 and 10 blows per inch. 

In this situation, the test record to select for signal matching analysis is readily 

apparent. An early, high energy blow, with good data quality should be 

selected and analyzed for the nominal resistance. The restrike blow count 

should be carefully recorded over the full restrike event as the rate by which 

the blow count decreases from inch to inch can be helpful. When the restrike 

blow count is less than 24 blows per foot, a lower energy restrike blow should 

be chosen for signal matching analysis to reduce the potential for 

overpredicting the nominal resistance.

In more difficult situations, limited pile movement may occur during restrike 

and several records may need to be analyzed with signal matching. 

Superposition of the activated shaft resistances under various restrike 

hammer blows may be used to assess the nominal pile resistance. In initial 

restrike blows, the shaft resistance may be mobilized along the upper portion 

of the pile shaft. Later restrike blows may indicate more shaft resistance on 

the lower portion of the pile once the upper shaft resistance has started to 

breakdown. The toe resistance and shaft resistance on the lower portion of 

the pile from the end of drive analysis should also be reviewed and, if 

appropriate, used in a superposition case. When using the toe resistance 

from an end of drive situation, the analyst should be confident that relaxation 

in the toe bearing layer is not a consideration or overestimation of the nominal 

resistance could result by using superposition.

• Difficulty to accept that, in some cases, dynamic measurements may provide

conservative predictions of the true geotechnical resistance and correlations and 

extrapolation between dynamic and static load test results are necessary.

Dynamic methods can yield conservative estimates of the true geotechnical 

resistance in some situations. Open ended pipe piles or H-piles which do not 

bear on rock may behave differently under dynamic and static loading 

conditions. Under dynamic loading conditions, the soil inside a pipe pile or 

between H-pile flanges may slip and produce internal shaft resistances. 
118

Under static loading conditions, this soil may plug and move with the pile

resulting in toe resistance over the full pile cross section. Hence both shaft 

and toe resistances may be different in open profile pile sections under static 

and dynamic loading conditions. Plugging behavior can also vary in different 

geomaterials. Careful interpretation and extrapolation of dynamic results is 

required in these situations.

• Understanding modeling uncertainties in signal matching analysis that can affect 

the reported soil resistance distribution.

 A portion of the soil resistance calculated on an individual soil segment in a 

signal matching analysis can usually be shifted up or down the shaft one soil 

segment without significantly altering the overall match quality. Similarly, it 

may be possible to shift a portion of the soil resistance from the last shaft 

segment to the pile toe or vis versa without significantly altering match quality. 

Therefore, use of the signal matching determined soil resistance distribution 

for uplift, scour, drag force, and other geotechnical considerations should be 

made with an understanding of these analysis limitations.
119

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Rausche, F., Moses, F., and Goble, G.G. (1972). Soil Resistance Predictions from 

Pile Dynamics. Journal of the Soil Mechanics and Foundations Division, 

ASCE, Vol. 98, No. 9, pp. 367-383.

Reiding, F.J., Middendorp, P., Schoenmaker, R.P., Middendorp, F.M. and Bielefeld, 

M.W. (1988). FPDS-2, A New Generation of Foundation Diagnostic 

Equipment, Proceeding of the 3rd International Conference on the Application 

of Stress Wave Theory to Piles, Ottawa, B.H. Fellenius, Editor, BiTech 

Publishers, pp. 123-134.

Verbeek, G. and Middendorp, P. (2011). Determination of Pile Damage in Concrete 

Piles. Deep Foundations Institute Journal, Vol. 6, No. 2, pp. 44-50.
122
123

CHAPTER 11

RAPID LOAD TESTING

The nominal resistance of driven pile foundations in axial compression can be 

evaluated by rapid load test methods. Rapid load tests methods can be applied to 

all driven pile types on land or over water. The test methods can provide time and 

cost savings where high loads are required or access is difficult. ASTM D7383,

Standard Test Methods for Axial Compressive Force Pulse (Rapid) Testing of Deep 

Foundations, provides additional details on rapid load test methods and procedures.

11.1 REQUIREMENTS FOR RAPID LOAD TESTS

In general, the previously stated reasons and prerequisites for a static load test 

program described in Section 9.1.1 are valid for rapid load test programs. Rapid 

load tests can be used to provide verification or refinement of the foundation design, 

confirm the nominal geotechnical resistance, and, in some instances, quantify the py response of laterally loaded piles. Knowledge of the subsurface stratigraphy 

including the engineering parameters of the geomaterials should be known prior to

performing a rapid load test.

Rapid Load Tests procedures are standardized in ASTM D7383 (2010). In this 

method, a loading apparatus generates a force pulse that will result in an applied 

pile head force versus time plot as shown in Figure 11-1. A target peak force is 

determined that should exceed the nominal geotechnical resistance plus the 

dynamic soil resistance. The target peak force is based on soil types, pile type and 

other project requirements. The applied force should exceed 50% of the actual peak 

force for a time duration of 4L/C or four times the pile length, L, divided by the pile 

material wave speed, C. The force applied must also exceed the static weight 

applied to the pile head due to the apparatus prior to the test, known as the pre-load 

force, for a time duration of at least 12L/C. Shorter time durations than 12L/C can 

be acceptable if additional force or movement measurement devices are used along 

the pile length at a distance described in the ASTM standard.

The force pulse should be measured by a calibrated force transducer placed 

between the loading apparatus and the pile head with the rated transducer capacity 
124

being at least 10% higher than the target peak force. The resulting pile head 

displacement should be measured by one or more calibrated displacement 

transducers. The ASTM standard also allows the Engineer to approve secondary 

pile head displacement measurement devices including redundant displacement 

transducers or accelerometers. As noted above, additional devices to measure 

force or displacement can be used along the pile’s length when the force pulse 

duration is less than 12L/C. All force and displacement measurements should be 

recorded versus time on an appropriate recording device with an appropriate 

sampling frequency and signal conditioning.

Figure 11-1 Typical axial compressive force pulse (after ASTM D7383).
125

11.2 BACKGROUND ON RAPID LOAD TEST METHODS

Rapid load testing can be performed using a combustion gas and a reaction mass or 

with a cushioned drop weight system. These methods will be described in 

subsequent sections.

11.2.1 Combustion Gas and Reaction Mass Apparatus (Statnamic)

The Statnamic testing method was developed in 1988 by Berminghammer 

Foundation Equipment and TNO, the Dutch governmental organization for applied 

scientific research. Bermingham and Janes (1989) described the method which 

uses solid fuel burned within a pressure chamber to rapidly accelerate upward the 

reaction mass positioned on top of the pile head. As the gas pressure increases, an 

upward force is exerted on the reaction mass, while an equal and opposite force 

pushes downward on the pile. Loading increases to a maximum and then unloads 

by a venting of the gas pressure. A load cell and accelerometers measure load and 

acceleration. Typically, the reaction mass weighs a minimum of 5% of the target 

peak force. The Statnamic test method is licensed to a single source in the US.

Statnamic tests for evaluation of static pile capacity have been performed on steel, 

concrete and timber piles. At present, individual piles, or pile groups with a 

combined static and dynamic resistance less than 9,000 kips can be tested. Axial 

compression tests have been conducted on both vertical and battered piles. The 

test method has been used on land and over water.

The principles of Statnamic can be described by Newton's Laws of Motion:

1. A body will continue in a state of rest or uniform motion unless compelled to 

change by an external force.

2. A body subjected to an external force accelerates in the direction of the 

external force and the acceleration is proportional to the force magnitude 

(F = ma).

3. For every action there is an opposite and equal reaction (F12 = -F21).

In the Statnamic test, a reaction mass is placed on top of the pile to be tested. The 

ignition and burning of the solid fuel creates a gas pressure force, F, that causes the 

reaction mass, m, to be propelled upward so that the acceleration amounts to about 
126

20 g's (F=ma). An equivalent downward force is applied to the foundation element, 

(F12 = -F21). The Statnamic concept is illustrated in Figure 11-2. 

Development began in 1988 with a Statnamic device capable of a 22 kip test load. 

From 1988 through 1992, the test load capability was incrementally increased to 

3600 kips. In 1994, a 6800 kip testing device was introduced. In 1998, a hydraulic 

catch mechanism was developed. The maximum test capacity was increased to 

9000 kips in 2005.

Figure 11-2 Statnamic concept (courtesy of Berminghammer Foundation 

Equipment). 

A base plate is attached to the pile head. The load cell, accelerometer, and piston 

base are positioned on top of the base plate. Next, the launching cylinder is placed 

on top of the piston base, thus enclosing the pressure chamber and propellant 

material. The segmental reaction mass is then stacked on the launching cylinder 

and a catching mechanism is placed around the reaction mass.

Depending upon the test load, a hydraulic catch, mechanical catch, or gravel 

retention structure is used to catch the reaction mass. The hydraulic catch system 

shown in Figure 11-3 is used for test loads of 1000 kips. A mechanical catch 

system is used for test loads of up to 4,400 kips and a gravel retention structure as 

shown in Figure 11-4 is used for loads of up to 9,000 kips. For the gravel retention 

structure, gravel backfill is placed in the annulus between the reaction mass and the 
127

retention structure. After propellant ignition and reaction mass launch, the granular 

backfill slumps into the remaining void to cushion the reaction mass fall.

The magnitude and duration of the applied load and the loading rate are controlled 

by the selection of piston and cylinder size, the fuel mass, the fuel type, the reaction 

mass, and the gas venting technique. The force applied to the pile is measured by 

the load cell. The acceleration of the pile head is monitored by the accelerometer 

and is integrated once to obtain pile head velocity and again to obtain displacement. 

Load and displacement data from the load cell and accelerometers are recorded, 

digitized, and displayed immediately in the field. Typical raw signals of the load and 

displacement records are given in Figure 11-5. These signals can then be 

converted into a raw load - displacement curve as given in Figure 11-6, which 

requires interpretation to derive the static pile capacity.

Figure 11-3 1000 kip hydraulic catch device on prestressed concrete pile

(courtesy of Applied Foundation Testing). 
128

Figure 11-4 Statnamic test in progress with gravel catch mechanism - 9,000 kip

device (courtesy of Applied Foundation Testing). 
129

Figure 11-5 Statnamic raw force and displacement measurements versus time 

(courtesy of Berminghammer Foundation Equipment). 

Figure 11-6 Statnamic load versus displacement

(courtesy of Berminghammer Foundation Equipment). 
130

11.2.2 Cushioned Drop Weight Systems

Rapid load tests can also be performed using cushioned drop weight systems that 

can generate the required peak force and force pulse duration. A drop weight, 

typically weighing between 5 and 15% of the target peak force, is mobilized to the 

site. A system (a crane, jack or some other mechanism) capable of lifting the drop 

weight to the required drop height and guiding it to strike the pile on center is also 

required. A load cell is once again located atop the pile head. However, in the 

cushioned drop weight systems, a spring system or a package of cushioning 

material is placed between the drop weight and the load cell. The springs or 

cushions lengthen the duration of the force pulse imparted to the pile. Some 

systems also include a clamping or catching mechanism to catch the rebounding 

drop mass after the force pulse has been applied. The catch mechanism prevents 

the application of additional force and improves the measurement of the pile head 

displacement from the main impact event.

The Pseudo Static Pile Load Tester, developed by Fundex (Schellingerhout and 

Revoort 1996), is a system with a clamp. This device, shown in Figure 11-7a, 

typically applies a series of drops from increasing release heights. Gradually, higher 

peak forces are applied and greater nominal resistances are mobilized, if present. 

The data acquisition and processing equipment is also shown in the photograph. 

Williams (2014) reported the system can achieve test loads up to 800 kips.

A large cushioned drop weight system for rapid load testing has also been reported 

by Miyasaka et al. (2009). In this system, named Hybridnamic, a modular ram of up 

to 85 tons is dropped on a cushion product specially developed to lengthen the force 

pulse duration. Other cushion materials, such as wood, are also viable. The drop 

height, drop weight and cushion thickness can be simulated with wave equation 

analysis. Drop weights of up to 170 kips are available, mobilizing nominal 

resistances of on the order of 1700 kips. A photograph of the Hybridnamic system is 

presented in Figure 11-7b. 

Another drop weight system that can be used for rapid load testing is the APPLE 

system (Rausche et al. 2008). This system, shown in Figure 11-8, can be 

configured with a ram weight of up to 160 kips. The system is capable of mobilizing 

nominal resistances of on the order of 1600 kips using rapid load test procedures.
131

Figure 11-7 Drop weight rapid load test systems: (a) Fundex system (courtesy 

Foundation Constructors Inc.) and (b) Hybridnamic system.

Figure 11-8 APPLE 32 ton modular rapid load test system with transducer.
132

11.3 RAPID LOAD TEST APPLICATIONS

Rapid load tests are primarily used to determine the nominal axial compressive 

resistance. A variety of pile types and sizes from small diameter timber, steel, and 

concrete piles up to large diameter open end steel pipes and concrete cylinder piles 

have been tested using rapid load tests. Internal and external strain gage 

instrumentation has also been used on rapid load tested piles for both nominal 

resistance interpretation as described in Section 11.4 and for load transfer 

assessments. 

Rapid load tests are attractive for testing large diameter open end pipe piles for two 

reasons. First, large diameter open end pipe piles often have high nominal 

resistances which are difficult to statically load test economically. Second, the 

longer duration, lower acceleration force pulse generated by a rapid load test, 

sometimes reduces slippage of the internal soil plug under the dynamic loading 

event.

Rapid load testing with a combustion gas apparatus has also been used for lateral 

load testing of piles and pile groups.

11.4 RAPID LOAD TEST INTERPRETATION METHODS

Initial correlations of Statnamic rapid load test results with static load tests for 

projects with toe bearing piles founded in till and rock showed good agreement 

without adjustment of the load - displacement results (Janes et al. 1991). However, 

in later tests, Statnamic rapid load test results overestimated the nominal resistance 

in some soils due to the dynamic loading rate effects (Janes and Campanella 1994). 

Several analysis procedures depending on pile length and pile response, as well as 

adjustment factors based on loading rate have subsequently been developed to 

derive the nominal resistance from Statnamic rapid load test results. These analysis 

procedures include the Unloading Point Method (UPM), the Modified Unloading 

Point Method (MUP), the Segmental Unloading Point Method (SUP), the Fully 

Mobilized UPM, and for cohesive soils the Sheffield Method. These methods are 

described in sections 11.4.1 through 11.4.5. Recommended loading rate reduction 

factors from NCHRP 21-08 by Paikowsky (2006) on Innovative Load Testing 

Systems for results analyzed with these rapid load test methods are discussed in 

Section 11.4.6. 
133

In addition to selecting the appropriate interpretation method and applying an 

appropriate loading rate reduction factor, rapid load tested piles should have 

sufficient maximum displacement of the pile head. Miyasaka et al. (2009) noted that 

rapid load tests using the unloading point method with a maximum displacement of 1 

to 3% of the equivalent pile diameter potentially overpredicted the nominal 

resistance compared to static load tests. Therefore, Miyasaka recommended 

running rapid load tests to a sufficiently large permanent displacement, on the order 

of 3% or more of the pile diameter. Holscher et al. (2012) recommended the 

maximum displacement of the pile head during the test be larger than 5% of the 

equivalent pile diameter to achieve geotechnical failure. Geotechnical failure is 

needed so that the dynamic soil resistance can be properly assessed and subtracted

from the nominal resistance.

Middendorp and Bielefeld (1995) proposed the wave number, Nw, as a guide for 

determining whether the Statnamic test was influenced by stress wave behavior and 

to determine the analysis procedure to be used. The wave number considers the 

foundation length, the wave speed of the pile material, and the duration of loading 

and is calculated from:

L

ct

L

D

N w L

w == Eq. 11-1 

Where:

Nw = wave number (unit less).

Dw = wave length (feet).

L = total pile length (feet).

c = wave speed of pile material (ft/s). 

tL = load duration (seconds).

11.4.1 Unloading Point Method (UPM)

The first widely used analysis method to adjust the raw Statnamic load - 

displacement results for dynamic loading rate effects was the Unloading Point 

Method (UPM) proposed by Middendorp et al. (1992).

In the UPM, all elements of the pile are assumed to move in the same direction and 

with almost the same velocity. According to the developers, this allows the pile to be 

treated as a rigid body undergoing translation as long as the pile has a wave 

number, Nw, greater than 12. The forces acting on the pile during a Statnamic test 

include the Statnamic induced load, Fstn, the pile inertia force, Fa, and the soil 

resistance forces which include the static soil resistance, Fu, the dynamic soil 
134

resistance, Fv, and the resistance from pore water pressure, Fp. A free body 

diagram of the forces acting on a pile during a Statnamic test is presented in Figure 

11-9. The soil resistance forces shown in the free body diagram are distributed 

along the pile shaft as well as at the pile toe.

In mathematical terms, the force equilibrium on the pile may be described as follows:

tFtFtFtFtF )()()()()( stn a u v +++= p Eq. 11-2 

Where:

Fstn = Statnamic induced load (kips).

Fa = pile inertia force (kips).

Fu = static soil resistance (kips).

Fv = dynamic soil resistance (kips). 

Fp = resistance from pore pressure (kips).

This equation may be rewritten in terms of static soil resistance as follows:

tFtFtFtFtF )()()()()( u stn a v −−−= p Eq. 11-3 

A simplifying assumption is made that the pore water pressure resistance, Fp, can be 

treated as part of the dynamic resistance, Fv. This simplifies the above equation to:

tFtFtFtF )()()()( u stn a −−= v Eq. 11-4 

Consider the Statnamic load - displacement data presented in Figure 11-10. The 

Statnamic load - displacement data can be separated into five stages. Stage 1 

includes the assembling of the Statnamic piston and reaction mass and thus is a 

static loading phase. The reaction mass is launched and Stage 2 therefore provides 

the initial loading of the dynamic event. The soil resistance is treated as linearly 

elastic. Pile acceleration and velocity are small, resulting in low inertia and damping 

forces on the pile.

Stage 3 is the basic load application portion of the cycle with fuel burning and 

pressure in the combustion chamber. In Stage 3, significant nonlinear soil behavior 

occurs as the pile and soil experience high acceleration and velocity. Thus the 

highest inertia and damping forces are generated in this stage. The maximum 

Statnamic applied load is reached at the end of Stage 3.
135

Figure 11-9 Free body diagram of pile forces in a Statnamic test

(after Middendorp et al. 1992).

Figure 11-10 Five stages of a Statnamic test (after Middendorp et al. 1992). 
136

In Stage 4, pressure in the combustion chamber is allowed to vent. Pile downward 

velocity and displacement continue but decrease throughout Stage 4. While the 

maximum Statnamic load is reached at the end of Stage 3, the maximum 

displacement occurs at the end of Stage 4. This is often due to the pile inertia force 

or significant dynamic resistance forces, Fv(t), but may also occur in soils with strain 

softening (the residual soil resistance is significantly lower than the peak resistance). 

Since the pile velocity is zero at the point of maximum displacement, tumax, the 

viscous damping, Fv(t), on the pile is also zero at the end of Stage 4 and the static 

pile capacity may be expressed only at that time as:

)()()( uu max stn u max ua max = − tFtFtF Eq. 11-5 

In Stage 5, the soil rebounds from the loading event and to achieve final equilibrium 

the pile unloads and rebounds as load and movement cease. The displacement at 

the end of Stage 5 is the permanent displacement or set experienced under the test 

event.

The data processing system records the applied Statnamic load and pile head 

acceleration and displacement throughout the test. The nominal static soil 

resistance, Fu, can then be calculated from the Statnamic load at the point of 

maximum displacement, Fstn(tumax), minus the pile inertia force. This nominal static 

soil resistance yields one point on the derived static load - displacement curve and 

may occur at a large displacement. If a limiting movement criterion such as 

described in Chapter 9 is used for load test interpretation, the nominal resistance 

may be less than this nominal static soil resistance.

To obtain the remaining points on the derived static load - displacement curve, the 

damping resistance, Fv, at other load - displacement points must be determined. 

Assuming all damping is viscous (e.g. linear), then the damping resistance force can 

be expressed in terms of a damping constant, C4, times the pile velocity at the 

corresponding time, v(t). The pile velocity is obtained by differentiating the 

measured pile head displacement.

If the maximum applied Statnamic load is greater than the nominal resistance, then 

the soil resistance at the beginning of Stage 4 through the point of maximum 

displacement at the end of Stage 4 will be a constant and will be equal to Fu(tmax), 

assuming the soil is perfectly plastic and does not exhibit strain hardening. The 

damping constant, C4, may be calculated from the maximum Statnamic load at the 

beginning of Stage 4, t4. This may be expressed as:
137

)(

)()()(

4

4 max 4

4 tV

tFtF ma t

C stn − uu − = Eq. 11-6 

Where:

C4 = damping constant.

Fstn = Statnamic induced load.

Fu = static soil resistance.

m = mass of pile. 

a = acceleration of pile. 

V = velocity of pile.

t4 = time at beginning of Stage 4 (see Figure 11-10).

tumax = time at maximum displacement (see Figure 11-10).

Assuming the damping constant, C4, is constant throughout the Statnamic loading 

event, the derived static load may be calculated at any point in time from:

)()()()( 4 tFtF ma tVCt u stn −−= Eq. 11-7 

Where:

Fu = static soil resistance.

Fstn = Statnamic induced load.

m = mass of pile. 

a = acceleration of pile. 

C4 = damping constant.

V = velocity of pile.

The derived Statnamic load - displacement curve is then constructed using the 

above equation and corresponding pile head displacement. An example of the 

derived load-displacement curve with the Unloading Point Method illustrating how 

the dynamic rate effects are subtracted from the Statnamic results is presented in 

Figure 11-11. 
138

Figure 11-11 Derived Statnamic load displacement curve with rate effects

(courtesy of Berminghammer Foundation Equipment).

11.4.2 Modified Unloading Point Method (MUP)

The Unloading Point Method rigid body assumption is not applicable for piles with a 

high toe resistance. On these piles, the pile head response (acceleration, velocity, 

and displacement) is significantly different than that at the pile toe. Because of the 

shortcomings of the UPM in this condition, the Modified Unloading Point Method 

(MUP) was developed by Justason (1997). The MUP method requires adding an 

additional accelerometer at the pile toe to define the toe behavior. The MUP method 

still assumes the pile to be a single mass but the acceleration of the mass is defined 

from the average of the pile head and toe displacement. The MUP method then 

uses the previously described UPM analysis procedure using the applied Statnamic 

force and the average accelerations and velocities.

11.4.3 Segmental Unloading Point Method (SUP)

Analytical studies by Brown (1995) have shown that the rigid body assumption used 

in the Unloading Point Method can result in overprediction of the nominal resistance 

and is not appropriate for long slender piles. Analysis of relatively long piles with a 

wave number, Nw, less than 10 was also problematic with the averaging techniques 
139

used in the Modified Unloading Point Method because of the time delay between the 

movement of the pile head and the movement of the pile toe and the resulting phase 

shift of the signals. To address this condition, the Segmental Unloading Point (SUP) 

Method was developed, and is described in Mullins et al. (2002). The SUP method 

separates the pile into discrete segments of shorter length where strain gages and 

accelerometers are used to calculate the force, acceleration, velocity and 

displacement of each segment. Details of the computation procedures may be 

found in Mullins et al. (2002), and in NCHRP 21-08, Paikowsky (2006).

The maximum number of segments is generally controlled by the number of strain 

gages. However, each strain gage level does not constitute a segment. Strain gage 

placement is usually determined by soil stratigraphy considerations. Multiple strain 

gages can be placed in a segment. The segment length is selected independent of 

gage location and must produce a wave number greater than 12.

The SUP method performs MUP analyses for each segment. The pile head derived 

static resistance is then calculated by summing the derived static response of each 

pile segment.

11.4.4 Fully Mobilized UPM

Miyasaka et al. (2009) proposed a modification to the unloading point method to 

avoid overprediction of the nominal resistance in cases where limited axial 

movement occurs. A case study was presented for a statically and rapidly load 

tested steel sheet pile installed in a mixed profile consisting of 9.5 feet of medium 

sand (SPT N=10), 8.9 feet of very loose silt (N=0), 16.4 feet of gravelly sand (N = 10 

to 20), and 3.9 feet of stiff, gravelly clay, underlain by fine dense sand (N = 50+). 

UPM results indicated an overprediction by as much as 1.43 times the nominal 

resistance if sufficient maximum and net pile head movement were not achieved. A 

comparison of the UPM results for each load cycle with the static load test result is 

presented in Figure 11-12. The drop height for rapid load tests range from 0.6 feet 

to 4.9 feet. Note the UPM and static load test results agree reasonable well once a 

maximum nominalized displacement exceeding 2.8% and a normalized net 

displacement in excess of 1.5% are achieved. Thus, for full mobilization of the 

geotechnical resistance using UPM results, Miyasaka proposed obtaining a 

minimum net permanent displacement of at least 3% of the pile diameter. ASTM 

D7383 also cautions on possible overprediction of the nominal resistance in rapid 

load test analysis cases with insufficient axial movement.
140

Figure 11-12 Normalized resistance versus normalized displacement 

(after Miyasaka et al. 2009).

11.4.5 Sheffield Method for Cohesive Soils 

Brown (2004), as well as Brown and Hyde (2006), describe an analysis method that 

incorporates loading rate effects in cohesive soils. Details of this method, the 

Sheffield Method, can be found in Holscher et al. (2012) Appendix D. The Sheffield 

Method incorporates the rate effect variation with pile settlement, and does not 

include rate reduction factors from Paikowsky (2006). It is applicable for piles 

installed in clay and assumes the majority of the nominal resistance is derived from 

shaft resistance. Holscher et al. (2012) cautioned strongly that this method not be 

used in sands or rock. The Sheffield Method depends on specific rate parameters, 

denoted α and β. It is preferable that the parameters α and β be determined from 

triaxial tests performed at variable displacement rates. However, if laboratory test 

data is not available, published literature values may be used. Rule of thumb values 

for α have also been empirically correlated to soil index tests from a relatively small 

data set, predominantly from cohesive soils from Europe, and β = 0.2. The Sheffield 

Method is show in Equation 11-8.

 

Eq. 11-8 
141

In which:

 Eq. 11-9 

Where:

Fu = static soil resistance.

Ft = force measured on pile head.

Fmax = maximum measured force applied during testing.

m = mass of pile.

a = acceleration of pile.

V = velocity of pile.

Vstatic= velocity used to determine parameters αα and ββ. 

Vo = normalizing constant.

α = material dependent parameter 1.

β = material dependent parameter 2.

PI = Plasticity Index.

11.4.6 Loading Rate Reduction Factors

In NCHRP 21-08, Innovative load Testing Systems, Paikowsky (2006) reported the 

correlation of 34 Statnamic rapid load test results with static load test results. The 

correlation database included driven H-piles, pipe piles and concrete piles as well as 

drilled shafts. Based on the correlation results, a loading rate reduction factor was 

recommended depending on the site soil or rock conditions. The loading rate 

reduction factor is to be applied to the derived static load-movement curve to 

account for overpredictions of the nominal resistance. Loading rate reduction factors 

of 0.96, 0.91, 0.69, and 0.65 were recommended for rock, sand, silt, and clay, 

respectively. Paikowsky recommended these loading rate reduction factors be 

applied to UPM, MUP, and SUP analyses of Statnamic test results.

Weaver and Rollins (2010) reviewed a different limited database of static load test 

and rapid load test results for drilled shafts with cohesive soils along the shaft and 

beneath the toe. That review indicated an average loading rate reduction factor of 

0.47 with a range of plus or minus 0.14 when used with the UPM. While not 

measured on driven piles, the results suggest a lower loading rate reduction factor 

should be considered when evaluating rapid load test results in cohesive deposits. 

This is particularly true if the pile head displacement is less than the 5% of the 

equivalent pile diameter needed to achieve geotechnical failure.
142

Brown and Powell (2013) reviewed the correlation between published static load 

tests and rapid load test results evaluated by UPM analysis. Based on this review, 

they proposed the loading rate reduction factor, ξ, for UPM analysis be selected 

based on the liquid limit, LL, of the clay soil in accordance with Equation 11-10.

 Eq. 11-10

Where: LL = Liquid Limit (%) 

Their compilation of correlation results is presented in Figure 11-13. 

Figure 11-13 Variation of the UPM loading rate reduction factor versus the soil liquid 

limit (after Brown and Powell 2013).

Correlations of rapid load test results with cushioned drop weight systems using the 

UPM, MUP or SUP analysis procedure with static load test results on driven piles in 

clays have not been reported in the literature. However, UPM, MUP or SUP 

analyses from cushioned drop weight systems would be subject to similar loading 

rate considerations.
143

11.4.7 Resistance Factors

AASHTO (2014) does not address resistance factors for rapid load test results 

analyzed with any of the previously described interpretation methods. Some 

agencies have stated resistance factor guidance or have sponsored research efforts 

to determined resistance factors for nominal resistance determined by rapid load test 

methods.

For redundant piles, the South Carolina Department of Transportation’s 

Geotechnical Design Manual, Version 1.1 (2010) suggests strength limit state 

resistance factors of 0.70 for a rapid load test result on a toe bearing pile in rock or 

very dense sand, and 0.65 for the nominal resistance determined by a rapid load 

test on a friction pile. For non-redundant piles, a reduced resistance factor of 0.55 is 

recommended for either end bearing or friction piles. These resistance factors for 

redundant and non-redundant piles are applicable to piles tested in axial 

compression with either the Statnamic rapid load test method or by high strain 

dynamic monitoring. The SCDOT manual notes that no increase in resistance factor 

is allowed on sites with multiple Statnamic tests unless the Statnamic rapid load test 

results are calibrated to a static load test.

A Florida Department of Transportation sponsored research project for 

determination of the resistance factor for rapid load tests was performed by McVay 

et al. (2003). The developed database was described as “small for statistical 

analysis purposes” and McVay calculated resistance factors using the FOSM 

approach to a reliability index, β, of 2.5. The calibration was performed using only 

the UPM analysis method with the loading rate reduction factors proposed by Mullins 

(2002), and summarized by Paikowsky (2006). The database used by McVay 

included 34 driven pile cases. Fifteen of the cases consisted of steel pipe piles 

ranging from 36 to 126 feet in length and 13 to 31 inches in diameter. Another 15 of 

the driven pile cases consisted of prestressed concrete piles ranging from 23 to 177 

feet in length and 16 to 36 inches in diameter. The data sets were from project sites 

in the United States, Canada, and Japan.

Due to the limited size of the correlation database, McVay recommended a 

resistance factor of 0.70 for the nominal axial resistance determined by a rapid load 

test on redundant driven piles in rock and non-cohesive soils, and a resistance factor 

of 0.60 for redundant driven piles in sands-clays-rocks mixed layers. For nonredundant driven piles, McVay recommended resistance factors of 0.60 and 0.50 for 

these soils conditions, respectively. For driven piles embedded primarily in clays, 

the Statnamic rapid load test method was not recommended by McVay without a 
144

calibrated static load test. This reflects the variability of the loading rate reduction 

factors described in the previous section. It should also be noted these resistance 

factors were developed using the unloading point method (UPM) of analysis, and 

that modified or segmental unloading point methods were not used in the calibration.

Cushioned drop mass apparatus testing has very little written in terms of rapid load 

tests calibrated to static load tests. Presumably, if signal matching is used on the 

rapid load test results as calibrated for high strain dynamic testing, the resistance 

factor for high strain dynamic testing with signal matching could be used. In all 

cases, the designer must decide on appropriate resistance factors considering 

redundancy of the foundation elements, number of tests, and the required reliability 

index prior to proceeding with a rapid load test program.

11.5 LATERAL LOADING APPLICATION

The use of the Statnamic test for lateral load application was also studied in NCHRP 

Report 461, Static and Dynamic Lateral Loading of Pile Groups by Brown et al. 

(2001). A lateral Statnamic test on a nine pile group is shown in Figure 11-14. The 

maximum lateral load applied to date in a Statnamic test is 2,700 kips. However, 

this is not a limit of the Statnamic test device but rather of the pile group response.

Figure 11-14 Lateral Statnamic test on nine pile group

(courtesy of Utah State University).
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11.6 CASE HISTORY

In 2004, a Statnamic rapid load test was conducted on a 42 inch O.D. x 0.75 inch 

thick wall, open end pipe pile. This 111 foot long pile was driven into a mixed soil 

profile. The soil conditions at the site were generally described as 46 feet of clay 

with interbedded layers of sands and silts over the upper 46 feet. This layer was 

underlain by sands to a depth of 95 feet, and then interbedded layers of sands and 

clayey silts from 95 to 118 feet. The rapid load test was performed on the open end 

pipe pile approximately 7 months after installation. The load-displacement plot for 

this test is presented in Figure 11-15.

Figure 11-15 Statnamic test result (courtesy Minnesota DOT).

The Statnamic test apparatus had a maximum capacity of 4270 kips and applied a 

load of 3840 kips. The Modified Unloading Point (MUP) Method was used to 

evaluate the Statnamic test result. The pile was originally driven without anticipating 

that a Statnamic test would be later conducted. Therefore, the pile was not 

equipped with a pile toe accelerometer. The testing firm assumed that the pile toe 

acceleration was one half of the measured pile head acceleration, and the average 

of the measured and assumed acceleration was then used to conduct the MUP 
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analysis. A loading rate reduction factor of 0.856 (weighted for the site stratigraphy) 

was also applied to the test result. The Statnamic derived static load-displacement 

curve was then evaluated according to the FHWA recommended static load test 

interpretation criterion for large piles in use at that time. That criterion defined failure 

as the sum of the elastic deflection plus the ratio of the pile diameter over 30. The 

assigned failure load following this approach was 2,480 kips.

11.7 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS

Advantages of rapid load testing relative to static testing include lower cost, shorter 

time required for test setup, and on-site mobility. The cost of a rapid load test can 

be on the order of one quarter to one half the cost of an equivalent static load test. 

The range in the cost of a rapid load test relative to a static load test depends upon 

the magnitude of the required load, the location of the site, and labor costs.

Cost savings may increase for rapid load tests when multiple tests are performed at 

the same site, or when rapid load tests are performed on piles having higher nominal 

resistances. Once a rapid load test apparatus is mobilized to a site, one or two rapid 

load tests can typically be performed in one day using the gravel catch structure. A 

greater number of rapid load tests per day can typically be performed using either 

the hydraulic catch device with the combustion gas and reaction mass apparatus or 

the cushioned drop weight systems.

The segmental reaction mass design for rapid load testing systems is advantageous 

as it allows assembly with relatively small hoisting equipment. In addition, since the 

reaction mass is typically 5 to 10 percent of the applied load, movement around a 

site for multiple tests is easier than for a static test.

Another advantage of a rapid load test is the mobilization of the soil resistance due 

to soil plugging in open ended sections. The slower loading rate in a rapid load test 

can be helpful in maintaining plugged behavior compared to a high strain dynamic 

test where the plug may slip under higher accelerations. Hence, plug behavior in a 

rapid load test may more closely resemble soil plug response in a static load test.

Rapid load tests do not require reaction piles. The smaller footprint associated with 

a rapid load test can be an advantage on sites with limited work areas.
147

Rapid load tests also have disadvantages. Rapid load test results must include both 

the influence of and correction for loading rate effects in all soil types. Correlations 

with conventional static tests are still being obtained to further address this issue as 

well as for calibration of rapid load test resistance factors.

The applied force pulse in a rapid load test necessary to mobilize the nominal 

resistance must be larger than the combined static and dynamic soil resistances. 

This disadvantage can be problematic in soils with high dynamic resistances. The 

pile in the presented rapid load test case history was driven into a mixed soil profile 

with low dynamic resistance based on the applied loading rate reduction factor of 

0.856. The Statnamic test applied load in this soil was 55% greater than the derived 

nominal resistance. The Statnamic applied pile stress was 40.3 ksi, or roughly 90% 

of the guaranteed yield strength of the pile material. The applied stress would have 

been larger to achieve this same nominal resistance in a soil profile having higher 

dynamic resistance.

A maximum net pile head displacement of at least 3% of the pile diameter is 

required for full mobilization of the soil resistance with the UPM analysis technique to 

obtain geotechnical failure. Miyasaka et al. (2009) demonstrated that normalized 

pile head displacements less than 3% can result in overprediction of the nominal 

resistance. 

Due to the limited number of rapid load tests systems and their locations, the 

mobilization cost for a rapid load test can be high. This is particularly true if the 

project requires only a single rapid load test, or if the test equipment must be 

mobilized from a distant location.

A limitation of a rapid load test can be instrumentation redundancy. Rapid load tests 

conducted without pile strain gage information lack the redundant check available in 

a conventional static load test (load cell and pressure gage) to check the calibration 

accuracy of the applied load.

11.8 PRACTICAL ISSUES AND CONSIDERATIONS

In a conventional static load test, a conservative estimate of the nominal 

geotechnical resistance is obtained when limited pile head movement occurs under 

the applied static loads. However, this is not necessarily the case in a rapid load 

test. When a rapid load test exhibits load-displacement behavior with a limited 

maximum displacement and permanent set, the nominal resistance may be either 
148

underpredicted or overpredicted. For piles driven into a dense granular deposit or 

on hard rock, a pile exhibiting a small permanent set likely indicates the nominal 

resistance has been underpredicted. However, in other soils and soft rock, the 

nominal resistance may be overpredicted as the force pulse is insufficient to cause 

geotechnical failure of the soil or soft rock and allow separation of the static and 

dynamic soil resistance effects. Rapid load tests in soils or soft rock should 

therefore have a permanent displacement of at least 3% of the pile diameter.

In layered soil profiles, a weighted loading rate reduction factor, ξ, must be 

determined and applied. The weighted loading rate reduction factor should be 

estimated based on the percentage of nominal resistance carried in a given layer 

and the appropriate loading rate reduction factor for the corresponding soil. Using 

the NCHRP 21-08 loading rate reduction factors described in Section 11.4.6, the 

loading rate reduction factor for a pile obtaining 50% of its nominal resistance in 

sand, ξ = 0.91, and 50% in clay, ξ = 0.65, would be 0.78. Similarly if 75% of the 

nominal resistance is obtained from the clay layer and 25% from the sand layer, the 

loading rate reduction factor would be 0.71.

In soils with high dynamic resistances, pile stresses should also be considered when 

selecting a rapid load test for nominal resistance verification. Pile stresses should be 

kept below the driving stress limits of the strength limit state when performing a rapid 

load test.
149

REFERENCES

ASTM D7383-10 (2010). Standard Test Methods for Axial Compressive Force Pulse 

(Rapid) Testing of Deep Foundations. Annual Book of ASTM Standards, Vol. 

4.08, ASTM International, West Conshohocken, PA, 9 p.

Bermingham, P. and Janes, M. (1989). An Innovative Approach to Load Testing of 

High Capacity Piles. Proceedings of the International Conference on Piling 

and Deep Foundations, London, England, Rotterdam Publishers, Vol. 1, 

pp. 409-413.

Brown, D.A. (1995). Closure – Evaluation of Static Capacity of Deep Foundations 

from Statnamic Testing. American Society of Testing and Materials (ASTM), 

Geotechnical Testing Journal, Vol. 18, No. 4, pp. 495-498.

Brown, D.A., O’Neill, M.W., Hoit, M., McVay, M., El Naggar, M.H., and Chakraborty, 

S. (2001). Static and Dynamic Lateral Loading of Pile Groups, NCHRP 

Report 461. Transportation Research Board – National Research Council, 

Washington, D.C., 50 p.

Brown, M.J. (2004). The Rapid Load Testing of Piles in Fine Grained Soils, Ph.D. 

Thesis, University of Sheffield, Sheffield, UK, 274 p.

Brown, M.J., Hyde, A.F.L., and Anderson, W.F. (2006). Analysis of a Rapid Load 

Test on an Instrumented Bored Pile in Clay, Geotechnique, Vol 56, Issue 9, 

pp. 627-638.

Brown M.J. and Powell, J.J.M. (2013). Comparison of Rapid Load Test Analysis 

Techniques in Clay Soils, Journal of Geotechnical and Geoenvironmental 

Engineering, Vol 139, No. 1, American Society of Civil Engineers, pp. 152-

161.

Holscher, P., Brassinga, H., Brown, M., Middendorp, P., Profittlich, M., and Tol, F. 

(2012). Rapid Load Testing on Piles – Interpretation Guidelines, CRC Press/ 

Baalkema, Leiden, 104 p.
150

Janes, M., Sy, A. and Campanella, R.G. (1994). A Comparison of Statnamic and 

Static Load Tests on Steel Pipe Piles in the Fraser Delta. Deep Foundations. 

Proceedings of the 8th Annual Vancouver Geotechnical Society Symposium, 

Vancouver, Canada, pp. 1-17.

Justason, M.D. (1997). Report of Load Testing at the Taipei Municiapl Incinerator 

Expansion Project, Taipei City, Taiwan.

McVay, M.C., Kuo, C.L., and Guisinger, A.L., (2003). Calibrating Resistance Factor 

in the Load and Resistance Factor Design of Statnamic Loading Test, 

University of Florida Project No. 4910450482312, Final Report, 129 p.

Middendorp, P., Bermingham, P. and Kuiper, B. (1992). Statnamic Load Testing of

Foundation Piles. Proceedings of the Fourth International Conference on the 

Application of Stresswave Theory to Piles, Balkema Publishers, A.A., The 

Hague, The Netherlands, pp. 581-588.

Middendorp, P., and Bielefeld, M.W. (1995). Statnamic Load Testing and the 

Influence of Stress Wave Phenomena, Proceedings of the First International 

Statnamic Seminar, Vancouver, Canada, pp. 207-220.

Mullins, G., Lewis, C., and Justason, M. (2002). Advancements in Statnamic Data 

Regression Techniques. Deep Foundations 2002: An International 

Perspective on Theory, Design, Construction, and Performance, American 

Society of Civil Engineers (ASCE), Geo Institute, GSP No.116, Vol. 2, pp. 

915-930.

Miyasaka T., Likins, G.E., Kuwabara, F., Rausche, F., and Hyodo, M. (2009). 

Improved Methods for Rapid Load Tests of Deep Foundations, Contemporary 

Topics in Deep Foundations, GSP 185, pp. 629-636.

Paikowsky, S. (2006). Innovative Load Testing Systems, NCHRP 21-08. National 

Cooperative Highway Research Program, Transportation Research Program, 

Washington, D.C., 148 p.

Rausche, F., Likins, G., Miyasaka, T., and Bulluck, P. (2008). The Effect of Ram 

Mass on Pile Stresses and Pile Penetration, Proceedings of the 8th

International Conference on the Application of Stress Wave Theory to Piles, 

IOS Press, Lisbon, Portugal, pp. 389-394.
151

SCDOT Geotechnical Design Manual Version 1.1 (2010). South Carolina 

Department of Transportation, Chapter 9, 7 p. 

Schellingerhout, A.J.G., and Revoort, E., (1996). Pseudo Static Pile Load Tester, 

Proceeding of the 5th International Conference on the Application of StressWave Theory to Piles, University of Florida, Gainesville, FL, pp. 1031-1037.

Weaver T.J., and Rollins, K.M. (2010). Reduction Factor for the Unloading Point 

Method at Clay Sites, Journal of Geotechnical and Geoenvironmental 

Engineering, Vol 136, No. 4, American Society of Civil Engineers, pp. 643-

646.

Williams, B. (2014). Rapid Load Testing course notes, Pile Load Testing Options 

Course, Pile Driving Contractors Association, Sacramento, CA.
152
153

CHAPTER 12

WAVE EQUATION ANALYSIS

12.1 WAVE EQUATION ANALYSIS INTRODUCTION

As discussed in AASHTO (2014) Article C10.7.3.8.5 as well as in Chapter 13 of this 

manual, dynamic formulas, together with observed penetration resistances, do not 

yield acceptably accurate predictions of actual nominal pile resistances. Moreover, 

they do not provide information on stresses in the piles during driving. The so-called 

“wave equation analysis” of pile driving has eliminated many shortcomings 

associated with dynamic formulas by realistically simulating the hammer impacts 

and pile penetration process. For most engineers, the term wave equation refers to 

a partial differential equation. However, for the foundation specialist, it means a 

complete approach to the mathematical representation of a system consisting of 

hammer, cushions, helmet, pile, and soil along with an associated computer 

program for the convenient calculation of the dynamic motions and forces in this 

system after ram impact.

The approach was developed by E.A.L. Smith (1960), and after the rationality of the 

approach had been recognized, several researchers developed a number of 

computer programs. For example, the Texas Department of Highways supported 

research at the Texas Transportation Institute (TTI) in an attempt to determine 

driving stresses and reduce concrete pile damage using a realistic analysis method. 

FHWA sponsored the development of both the TTI program (Hirsch et al. 1976) and 

the WEAP program (Goble and Rausche 1976). FHWA supported the WEAP 

development to obtain analysis results backed by measurements taken on 

construction piles during installation for a variety of hammer models. The WEAP 

program was updated several times under FHWA sponsorship, until 1986 (Goble 

and Rausche 1986). Later, additional options, improved data files, refined 

mathematical representations and modern user conveniences were added to this 

program on a proprietary basis, and the program is now known as GRLWEAP, (Pile 

Dynamics, Inc. 2010). Similar computer programs have been developed, such as 

PDPWAVE (Bielefeld and Middendorp 1992), that are based on the method of 

characteristics, a mathematical model that differs from Smith’s lumped mass 

model.
154

The wave equation approach has been subjected to a number of checks and 

correlation studies. Studies on WEAP performance of have produced publications 

demonstrating that program's performance and utility (e.g., Blendy 1979; Soares et 

al. 1984; Rausche et al. 2004). Documentation of the most recent version of this 

program, GRLWEAP 2010, has been prepared by Pile Dynamics, Inc. (2010).

AASHTO (2014) design specifications recommend a resistance factor of 0.50 when 

wave equation analysis is used for the determination of the nominal resistance 

without the benefit of dynamic measurements or load test results but with field 

confirmation of hammer performance. The wave equation database used for LRFD 

calibration consisted of nominal resistance predictions determined using the 

GRLWEAP program and its associated pile and soil models.

This chapter explains what a wave equation analysis is, how it works, and what 

problems it can solve. Example problems, highlighting wave equation program 

applications, are demonstrated. Basic wave equation program input and output are 

presented for demonstration purposes using the GRLWEAP software. However, this 

should not be construed as a promotion or endorsement of this particular software 

product by the FHWA.

12.2 WAVE PROPAGATION

Input preparation for wave equation analyses is often very simple, requiring only 

very basic driving system and pile parameters in addition to a few soil parameters for 

which standard recommendations are given. Thus, a wave equation program can 

be run with minimal specialized knowledge. However, interpretation of calculated 

results is facilitated, and errors in result application may be avoided, by knowledge 

of the mechanics of stress wave propagation and familiarity with the particular 

project’s design requirements and constraints.

In the first moment, after a hammer has struck the pile top, only the pile particles 

near the ram-pile interface are compressed. This compressed zone, or force pulse, 

as shown in Figure 12-1, expands into the pile toward the pile toe at a constant wave 

speed, C, which depends on the pile's elastic modulus and mass density (or specific 

weight). When the force pulse reaches the embedded portion of the pile, its 

amplitude is reduced by the action of static and dynamic soil resistance forces. 

Depending on the magnitude of the soil resistances along the pile shaft and at the 

pile toe, the force pulse will reflect from the pile toe either as a tension or a 

compression force pulse, which travels back to the pile head. Both incident and
155

Figure 12-1 Wave propagation in a pile (adapted from Cheney and Chassie 2000). 

reflected force pulses will cause a pile toe motion and produce a permanent pile set 

if their combined energy and force are sufficient to overcome the static and dynamic 

resistance effects of the soil.

12.3 WAVE EQUATION METHODOLOGY

In a Smith-type wave equation analysis, the hammer, helmet, and pile are modeled 

by a series of segments each consisting of a concentrated mass and a weightless 

spring. The hammer and pile segments are approximately one meter in length. 

Shorter segments occasionally improve the accuracy of the numerical solution at the 

expense of longer computer run times (Rausche et al. 2004). Spring stiffness and 

mass values are calculated from the cross sectional area, modulus of elasticity, and 

specific weight of the corresponding pile section. Hammer and pile cushions are 

represented by additional springs whose stiffness are calculated from area, modulus 

of elasticity, and thickness of the cushion materials. In addition, coefficients of 
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restitution (COR) are usually specified to model energy losses in cushion materials, 

and in all segments, which can separate from their neighboring segments by a 

certain slack distance. The COR is equal to 1.0 for a perfectly elastic collision which 

preserves all energy, and is equal to 0.0 for a perfectly plastic condition which loses 

all deformation energy. The usual condition of partially elastic collisions is modeled 

with an intermediate COR value. For example, the default value for the COR of 

wood is 0.50. However, when a softwood cushion is badly worn, a value of 0.25 

may be more appropriate and can be entered into the wave equation input together 

with other cushion properties.

The soil resistance along the embedded portion of the pile and at the pile toe is 

represented by both static and dynamic components. Therefore, both a static soil 

resistance force which is pile displacement related, and a dynamic soil resistance 

forces which is pile velocity related act on every embedded pile segment. The static 

soil resistance forces are modeled by elasto-plastic springs and the dynamic soil 

resistance by dashpots. The displacement at which the static soil resistance 

changes from elastic to plastic behavior is referred to as the soil "quake". In the 

Smith damping model, the dynamic soil resistance is proportional to a damping 

factor times the pile velocity times the temporary static soil resistance. Additional 

discussion on quake and damping factors is presented in Section 12.6.7. A 

schematic of the wave equation hammer-pile-soil model is presented in Figure 12-2. 

As the analysis commences, a calculated or assumed nominal resistance (ultimate 

capacity, Rut) from user specified values is distributed along the shaft and toe 

according to user input or program calculation among the elasto-plastic springs. 

Similarly, user specified damping factors are assigned to shaft and toe to represent 

the dynamic soil resistance. The analysis then proceeds by calculating a ram 

velocity based on hammer efficiency and stroke inputs. The ram movement causes 

displacements of helmet and pile head springs, and therefore compressions (or 

extensions) and related forces acting at the top and bottom of the segments. 

Furthermore, the movement of a pile segment causes both static and dynamic soil 

resistance forces. A summation of all forces acting on a segment, divided by its 

mass, yields the acceleration of the segment. The product of acceleration and time 

step summed over time is the segment velocity. The velocity multiplied by the time 

step yields a change of segment displacement which then results in new spring 

forces. These spring forces divided by the pile cross sectional area at the 

corresponding section equal the stress at that point.

Similar calculations are made for each segment until the accelerations, velocities 

and displacements of all segments have been calculated during the time step. The 
157

analysis then repeats for the next time step using the updated motion variables, 

velocity and displacement, of the segments from the previous time step. From this 

process, the accelerations, velocities, displacements, forces, and stresses of each 

segment are computed over time. Additional time steps are analyzed until the pile 

toe begins to rebound.

Figure 12-2 Typical wave equation pile and soil models.
158

The permanent set in inches of the pile toe is calculated by subtracting a weighted 

average of the shaft and toe quakes from the maximum pile toe displacement. An 

alternation calculation option available in the GRLWEAP program is called Residual 

Stress Analysis (RSA) which analyzes several hammer blows in sequence and then 

determines the permanent set from the pile top displacements of consecutive blows. 

However, to date, this more accurate approach has only been adopted for and 

checked against records from fluted and tapered piles. The inverse of the 

permanent set is the penetration resistance (blow count) in blows per foot or blows 

per inch that corresponds to the input nominal resistance. By performing wave 

equation analyses over a wide range of nominal resistances a curve or "bearing 

graph" can be plotted which relates nominal resistances to the pile penetration

resistance or blow count.

A wave equation bearing graph is substantially different from a similar graph 

generated from a dynamic formula. The wave equation bearing graph is associated 

with a single driving system, hammer stroke, pile type, soil profile, and a particular 

pile length. If any one of the above items is changed, the bearing graph will also 

change. Furthermore, wave equation bearing graphs also include the maximum 

calculated compression and tension stresses.

In addition to the bearing graph, wave equation analysis can provide two alternative 

results, the constant capacity analysis, or "inspector's chart", and the “drivability 

analysis.” The inspector's chart establishes a relationship between variable hammer 

energy or stroke and pile penetration resistance for one particular, user specified, 

nominal resistance (ultimate capacity) value. Associated stress maxima are also 

included in the chart, enabling the user to select a practical hammer energy or stroke 

range both for reasonable penetration resistances and driving stress control. This 

analysis option is described in greater detail in Section 12.5.2.

The drivability analysis calculates penetration resistances and stresses from user 

input shaft and toe resistance values at up to 100 user selected pile penetrations. 

The calculated results can then be plotted together with the nominal resistance 

values versus pile penetration. The resulting plot would depict those pile 

penetrations where refusal might be expected or where dangerously high driving 

stress levels could develop. In addition, a crude estimate of pure driving time (not 

counting interruptions) is provided by this analysis option. The drivability option is 

described in greater detail in Section 12.5.3.
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12.4 WAVE EQUATION APPLICATIONS

A bearing graph provides the wave equation analyst with two types of information:

1. It establishes a relationship between nominal resistance and pile penetration 

resistance or blow count. From the user's input data of the resistance values 

along shaft and at the toe, the wave equation analysis estimates the 

permanent set in inches under one hammer blow. Specifying up to ten 

nominal resistance (ultimate capacity) values yields a relationship between 

nominal resistance and penetration resistance in blows per foot or blows per 

inch.

2. The analysis also relates driving stresses in the pile to the pile penetration 

resistance.

3. The analysis also relates hammer stroke or hammer energy to the pile 

penetration resistance for a given nominal resistance.

The user typically develops a bearing graph or an inspector's chart for different pile 

lengths and uses these graphs in the field, with the observed penetration resistance, 

to determine when the pile has been driven to the required nominal resistance.

In the design stage, the foundation engineer should select typical pile types and 

driving equipment known to be locally available. Then by performing the wave 

equation analysis with various equipment and pile size combinations, it becomes 

possible to rationally:

1. Design the pile section for drivability to the required depth and/or nominal 

resistance.

For example, design considerations such as scour or consolidation settlements due 

to lower soft layers may make it necessary to drive a pile through a hard layer whose 

penetration resistance exceeds the resistance expected at final penetration. A thin 

walled pipe pile may have been initially chosen during design. However, when this 

section is checked for drivability, the wave equation analysis may indicate that even 

the largest hammers will not be able to drive the pipe pile to the required depth, 

because it is too flexible (its impedance is too low). Therefore, a wall thickness 

greater than necessary to carry the factored load has to be chosen for drivability 

considerations. (Switching to an H-pile or predrilling may be other alternatives).
160

2. Aid in the selection of pile material properties to be specified based on 

probable driving stresses in reaching penetration and/or nominal resistance 

requirements.

Suppose that it would be possible to drive a thinner walled pipe pile or lower weight 

H-pile section to the desired depth, but with excessive driving stresses. More 

cushioning or reduced hammer energy would lower the stresses, but would result in 

a refusal penetration resistance. Choosing a high strength steel grade for the pipe 

or H-section could solve this problem. For concrete piles, higher concrete strength 

and/or higher prestress levels may provide acceptable solutions.

3. Support the decision for a new penetration depth, factored load, and/or 

different number of piles.

In the above example, after it has been determined that the pile section or its 

material strength had to be increased to satisfy pile penetration requirements, it may 

have become feasible to increase the design load of each pile and to reduce the 

total number of piles. Obviously, these considerations would require revisiting 

geotechnical and/or structural considerations.

Once the project has reached the construction stage, additional wave equation 

analyses should be performed on the actual driving equipment by:

1. Construction engineers – for hammer approval and cushion design.

Once the pile type, material, and pile penetration requirements have been selected 

by the foundation designer, the hammer size and hammer type must be selected. 

These parameters may have a decisive influence on driving stresses. For example, 

a hammer with adjustable stroke or fuel pump setting may have the ability to drive a 

concrete pile through a hard layer while allowing for reduced stroke heights and 

increased tension stress control when penetrating soft soil layers.

Cushions are often chosen to reduce driving stresses. However, softer cushions 

absorb and dissipate greater amounts of energy thereby increasing the penetration 

resistance. Since it is both safer (reducing fatigue effects) and more economical to 

limit the number of blows applied to a pile, softer cushions cannot always be chosen 

to maintain acceptable driving stresses. Also, experience has shown that the 

addition of hammer cushion material is relatively ineffective for limiting driving 

stresses. 
161

Hammer size, energy setting, and cushion materials should always be chosen such 

that the maximum expected penetration resistance is less than 120 blows/ft. One 

exception to this upper limit is piles driven through soft soils that transition abruptly 

to hard rock. In this unique case, piles are often final seated on the hard rock at 5 

blows per 1⁄4 inch (240 blows/foot equivalent) and the use of a larger hammer may 

result in driving stress concerns.

For most situations where wave equation simulation is used for hammer selection, it 

is prudent to select a hammer that can achieve the required nominal resistance at a 

blow count not exceeding 100 blows/ft. The blow count should also be greater than 

30 blows/t for reasonably accurate construction control. This is required, because 

(a) the relative error of an inaccurate blow count measurement is greater for lower 

penetration resistances and (b) the dynamic methods of nominal resistance 

assessment tend to over-predict when driving is very easy. Of course, adjustable 

hammers may be accepted based on their lower energy settings. Exceptions should 

also be made when the accuracy of the blow count measurement is irrelevant. Such 

situations arise when the pile has to be driven to depth at expected capacities above 

the required minimum or because a large component of the nominal resistance is 

derived from soil setup.

2. Contractors – to select an economical driving system to minimize installation 

cost. 

While the construction engineer is interested in a safe pile installation method, 

contractors would like to minimize equipment size and driving time for cost 

considerations. Light weight, simple, and rugged hammers which have a high blow 

rate are obviously preferred. The wave equation analysis can be used to roughly 

estimate the anticipated number of hammer blows and the time of driving. This 

information is particularly useful for a relative evaluation of the economy of driving 

systems. 

For concrete piles, additional considerations might include the cost of pile cushions, 

which are usually discarded after a pile has been installed. Thus, thick plywood pile 

cushions may result in a considerable cost.

Near refusal penetration resistances are particularly time-consuming. Since it is 

known that stiffer piles drive faster with lower risk of damage, the contractor may

choose to upgrade the wall thickness of a pipe pile or the section of an H-pile for 

improved overall economy.
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12.5 WAVE EQUATION EXAMPLES

This section presents several examples that illustrate the application of the wave 

equation analysis for the solution of design and construction problems. The 

resistance factor applied to the nominal resistance in the following examples is 0.80 

unless otherwise stated in the example. This assumes that a static pile load test 

was performed in addition to and to calibrate the wave equation result. As noted in 

Chapter 7, a resistance factor of 0.65 should be applied to the nominal resistance if 

dynamic testing with signal matching would be performed on the project instead of a 

static load test. If only wave equation analyses were used for resistance verification, 

then a resistance factor of 0.50 would be applicable. Furthermore, the nominal 

resistance times the resistance factor in a wave equation analysis should be greater 

than the sum of the critical factored loads plus resistances from any overlying layers 

unsuitable for or not present during long term support. Note that the nominal 

resistance (Chapter 7) corresponds to ultimate capacity in the literature. Hence, 

input or output for wave equation analysis programs often identify the nominal 

resistance as ultimate capacity or Rn

12.5.1 Example 1 – General Bearing Graph

One of the primary applications of a wave equation analysis is to develop a bearing 

graph relating the nominal resistance to the pile penetration resistance or blow 

count. A bearing graph is often used to develop the driving criterion. For a desired 

nominal resistance, the required blow count can be obtained from the bearing graph. 

Consider the soil profile in Figure 12-3. In this example, the foundation designer has 

estimated that a 65 foot long, 14 inch by 0.312 inch wall, closed end pipe pile with a

steel yield strength of 35 ksi will be driven 62 feet into a deep deposit of medium 

sands. The foundation designer anticipates that the pile would achieve a nominal 

resistance of 350 kips at that depth. Using only wave equation analysis for the field 

verification method, the AASHTO resistance factor is 0.50. This would allow a 

maximum factored load of 0.50 times 350 kips or 175 kips. The designer also 

anticipated that the resistance on the shaft (pile toe) would be 52% (48%) of the total 

nominal resistance and that it would be triangularly distributed. Note that the 

calculated resistance is the expected long term nominal resistance and because the 

soil type is sand it can be expected that the resistance during driving will be 

essentially the same. Also, based on program recommendations for this 

displacement pile driven into a medium sand, the toe quake is input as the pile 

diameter divided by 60 (0.23 inches) while the shaft damping in these cohesionless 

soils is set to 0.05 s/ft. Shaft quake and toe damping are generally left at the 
163

defaults of 0.10 inches and 0.15 s/ft, respectively. These recommendations have, 

on the average, yielded satisfactory agreement with field observation and dynamic 

testing results. However, once piles have been driven at specific sites, load tests 

may indicate that different parameters may be more appropriate. Figure 12-3

summarizes the input quantities.

The contractor selected a Delmag D12-42 single acting diesel hammer for driving 

the pipe piles. The contractor's hammer submittal indicates that the hammer 

cushion will consist of 1 inch of aluminum and 1 inch of Conbest with a cross 

sectional area of 227 in2 and an averaged (Conbest and aluminum) elastic modulus 

of 530 ksi. A helmet weight of 1.7 kips is reported.

Figure 12-3 Example 1 – Problem profile. 

Based on this information, a wave equation analysis can be performed. The nominal 

resistance of 350 kips is input along with selected additional values for which the 

wave equation analysis calculates the net permanent displacement and, inversely, 

the penetration resistance in blows/foot. By plotting calculated penetration 

resistances versus the corresponding input nominal resistance values, a bearing 

graph is developed. The results of the example analysis are shown in Table 12-1

and in Figure 12-4. 
164

In the bottom half of the bearing graph, the nominal resistance versus penetration 

resistance in blows/foot is represented by the solid line. This graph shows that for a 

nominal resistance (ultimate capacity) of 350 kips, a penetration resistance of 99 

blows/foot is required. This is less than the AASHTO recommended limit of 120 

blows/foot. Although this is a relatively high blow count, it will neither require an 

excessive pile driving effort nor would it be too low to be inaccurate as far as 

construction control is concerned.

Table 12-1 Bearing Graph Numerical Results

Nominal 

Resistance 

 

kips 

Maximum 

Compression 

Stress 

ksi 

Maximum 

Tension 

Stress 

ksi 

Penetration 

Resistance 

(Blow Count) 

bl/ft 

Hammer

Stroke 

ft 

Transferred 

Energy 

ft-kips 

50 19.15 1.19 5.1 5.70 17.3

100 22.40 0.58 11.8 6.54 15.8

150 24.52 1.57 20.7 7.21 15.6

200 25.82 1.52 29.9 7.64 15.7

250 27.17 2.05 42.5 8.16 16.3

300 28.14 2.21 63.4 8.52 16.9

350 28.89 2.18 98.8 8.82 17.3

400 29.34 2.11 169.6 8.99 17.4

450 29.97 2.11 279.9 9.24 17.7

500 30.34 2.08 631.3 9.41 17.9 

Also in the bottom half of the bearing graph, the corresponding hammer stroke 

versus penetration resistance is represented by the dashed line. This curve is 

important for variable stroke hammers as a check on hammer performance when the 

driving criterion is applied. In this case, the penetration resistance of 99 blows/foot 

for 350 kip nominal resistance is based upon the hammer operating at a hammer 

stroke of 8.8 feet. Should field observations indicate significantly (say more than 

10% difference) higher or lower strokes, then a lower or higher penetration 

resistance would be necessary for the same nominal resistance, because the 

hammer force and/or energy would be higher or lower. Hammer stroke information 

is therefore essential for field evaluation and control of the pile installation process.

An inspector’s chart analysis (see Example 2) provides this information determining 

required blow count as a function of hammer stroke. For significantly differing 

hammer field performance, a new wave equation analyses would be necessary with 

a modified maximum combustion pressure or a fixed input stroke.
165

Figure 12-4 Typical wave equation bearing graph. Top plot – driving stresses; 

bottom plot – nominal resistance and diesel hammer stroke.
166

In the upper half of Figure 12-4, maximum compression and tension driving stresses 

are plotted as a function of penetration resistance. Of primary interest for a steel pile 

is the compression driving stress, which is represented by the solid black line. This 

plot shows that, at 99 blows/foot (and at the required nominal resistance), the 

maximum compression stress calculated in the pile is almost 29 ksi. This is less 

than 90% of the yield strength of 35 ksi or 31.5 ksi. This compression stress level is 

acceptable according to Article 10.7.8 of AASHTO (2014) design specifications and 

as discussed further in Chapter 8. However, any non-uniform stress components 

(e.g. bending caused by poor hammer-pile alignment or pile-toe contact with sloping 

rock) are not included in the wave equation results and would be additional. In any 

case, the 90% yield limit applies to the stress averaged over the pile cross section.

Though the analysis was conducted for an estimated penetration of 62 feet, in the 

field the required penetration resistance may be reached at a lesser or greater 

depth. The static analysis only serves as an initial estimate of the required 

penetration depth. The actual driving behavior and construction monitoring will 

confirm whether or not the static calculation was adequate. If the actual driving 

behavior is significantly different from the analyzed situation (say the required blow 

count is already reached at 50 feet of penetration), an additional analysis should be 

performed to better match field observations. In general, the nominal resistance 

versus pile penetration resistance relationship is relatively insensitive to changes in 

the penetration depth and, therefore, to the distribution of the resistance along the 

pile unless there is a significant change in the soil profile. Of course, if unexpected 

changes in hammer performance, driving system components, or soil properties 

appear to cause the difference, then it would be prudent to check the equipment 

performance and soil resistance by dynamic measurements. Higher penetration 

resistances from penetrating embankment fills or scour susceptible material, etc., 

should also be considered in this assessment.
167

12.5.2 Example 2 – Constant Capacity / Variable Stroke Option

The hammer-pile-soil information used in Example 1 will be reused for a constant 

capacity (or inspector's chart) analysis in Example 2. In this example, the 

penetration resistance required for the 350-kip nominal resistance is evaluated over 

a range of hammer strokes. The resulting inspector’s chart would be helpful for field 

personnel in determining when pile driving can be terminated if the field observed 

hammer stroke varies from that originally predicted by the wave equation bearing 

graph analysis. Figure 12-5 shows the resulting inspector’s chart. The lower half of 

Figure 12-5 presents the hammer stroke versus penetration resistance plot for the 

required nominal resistance of 350 kips. Where the point of intersection of the 

observed stroke and penetration resistance plots below the curve, the nominal 

resistance has not been obtained. Any combination of stroke and penetration 

resistance plotting above the curve indicates that the required resistance level has 

been reached. For example, any stroke greater than 8.2 feet at a penetration 

resistance of 90 blows/foot is acceptable. The upper half of either graph shows the 

stress maxima associated with a particular driving resistance. Hence, the 

inspector’s chart analysis aids the inspection personnel in field control.

12.5.3 Example 3 – Drivability Studies

Scour and seismic design considerations often result in increased pile penetration 

requirements. Therefore, the ability of a given pile to be driven to the required 

penetration depth should be evaluated in the design stage in a wave equation 

drivability study, as presented in this example.

Figure 12-6 illustrates the installation conditions at an interior bridge pier in a river. 

A cofferdam will be required for pier construction. The interior of the cofferdam will 

be excavated 16.5 feet below riverbed prior to pile installation. The excavated layer 

consists of loose silt. Below the silt, a 13.5 foot thick layer of extremely dense sand 

and gravel layer with some clay was noted which in turn overlies a medium dense 

sand layer. Bedrock was encountered approximately 70 feet below riverbed.

The contractor has selected a Conmaco C 160 air hammer which has a ram weight 

of 16.25 kips, a rated stroke of 3.0 feet for a rated energy of 48.75 ft-kips. The 

hammer cushion consists of Nylon, the helmet weight is 4.07 kips and the pile 

cushion was chosen as 10 inches of plywood. A new pile cushion is to be used for 

every pile. For that reason the elastic modulus of 30 ksi for “new plywood” is used at 
168

Figure 12-5 Constant capacity analysis (inspector’s chart).
169

Figure 12-6 Example 3 – Problem profile.

the beginning of driving. However, as the plywood compresses during driving and 

gets harder, a “used plywood cushion” elastic modulus of 75 ksi should be chosen to 

reflect the associated higher stiffness of the harder cushion at end of driving. This 

recommendation is based on correlation studies (Rausche et al. 2004) which 

indicated that the stiffness of plywood cushions at the end of driving are typically 2.5 

times higher than in their new condition (for oak boards the factor is 1.5). Note that 

for the correctly effecting the stiffness increase, only the elastic modulus should be 

modified while the cushion thickness is always entered as the nominal thickness in 

the “new” cushion conditions. 

The extremely dense sand and gravel layer was estimated to have a soil friction 

angle φ = 43o. This friction angle was used in the static calculations of toe 

resistance. However, the friction angle was limited to 36o for the hard angular gravel 

when calculating shaft resistance. The friction angle for the medium dense sand 

layer was estimated to be φ = 33o

. 

During construction, the silt soils will be removed from within the cofferdam area. 

However, the silt soils outside the cofferdam will still be present at the time of 

construction. Therefore, the soil resistance to pile driving should be calculated with 

consideration for the overburden pressure from these materials. However, hydraulic 
170

experts predict that the 16.5 feet of loose silt may erode completely due to channel 

degradation scour. Thus, for long term pile nominal resistance, static calculations 

should ignore the effective weight of the silt layer. As a result, a higher soil 

resistance than required to meet the static load requirements must be anticipated 

during pile installation.

Initial static analysis indicates that a 14 inch square prestressed concrete pile would 

develop the required nominal resistance of 400 kips, primarily through end bearing, 

at a depth of 10 feet below the cofferdam excavation level (26.5 feet below river 

bed). However, when considering the reduction in the effective overburden pressure 

from the scouring of the silt layer, the pile would have an nominal resistance of only 

241 kips at a penetration depth of 10 feet below cofferdam excavation level and 318 

kips immediately above the dense sand where the end bearing would be unreliable. 

Additional static nominal resistance calculations were performed at increased pile 

penetration depths for the pre-scour profile. These analyses show that when 

punching through the upper, extremely dense sand layer the nominal resistance 

would at first be lower in the dense sand and gravel but would again reach a 400 kip 

nominal resistance at a depth of 49 feet below cofferdam excavation level (65.5 feet 

below river bed).

For the installation condition, a pre-scour analysis indicates that the nominal 

resistance at a depth of 49 feet would be 472 kips and this is the soil resistance that 

must be overcome at the end of driving. (An almost identical driving resistance 

exists at a depth of 13 feet depth below excavation (29.5 feet below river bed)).

Next, nominal unit resistance values for both shaft and toe resistance versus depth 

with consideration of the silt overburden stress were calculated by static analysis 

and input into a wave equation program. The soil profile consists primarily of sandy 

materials. Significant soil setup or relaxation in these materials is not considered 

likely and therefore no gain or loss factors due to driving had to be considered in the 

drivability analysis (gain/loss factors for both shaft and toe were set to 1.0). Since 

the study is conducted in the design stage, the use of a locally available single 

acting air hammer driving system was assumed.

The following additional input considerations should be mentioned:

1. The concrete pile is a displacement pile. For that reason, the toe quake for 

the extremely dense sand and gravel was set to the pile width, D, divided by 

120 (0.12 inches). 

2. For the dense sand the toe quake was set to D/60 – 0.23 inches.
171

3. The pile cushion becomes stiffer (elastic modulus increases and thickness 

decreases) during pile installation. GRLWEAP Help (Rausche et al. 2004), 

recommends using a pile cushion stiffness for the end of driving which is 2.5 

times the initial stiffness which is based on an elastic modulus of 30 ksi. This 

correction factor includes both the effect of the cushion reduction in thickness 

as well as the increase in elastic modulus. Thus for the various depths 

analyzed, the cushion stiffness was gradually increased between the initial 

and final driving depths.

The drivability analysis result (Figure 12-7) indicated that the 14 inch concrete pile 

would encounter a maximum penetration resistance of 158 blows/foot in the upper 

extremely dense sand and gravel deposit just before breaking through to the dense 

sand layer. For the final depth of 49 feet (65.5 feet below river bed) a penetration 

resistance of 60 blows/foot was calculated. Since the high blow count in the 

extremely dense sand layer is only present for a short distance, it could be 

concluded that the 14 inch concrete pile could be driven to the required penetration 

depth of 49 feet. This might be an erroneous conclusion. Although the static 

analysis would likely provide an adequate assessment of soil resistance for the first 

pile driven, an increase in the friction angle from group densification could 

significantly affect the resistance to driving of additional displacement piles, 

particularly within the tight confinement of the cofferdam. Also, dense deposits tend 

to develop negative pore pressures during shear, resulting in temporary increases in 

soil resistance which may later dissipate. If it is assumed that these factors cause a 

30% increase in both shaft and toe resistances during the driving of subsequent 

piles, a second drivability analysis for the densified condition would indicate that the 

later piles essentially refuse (blow count greater than 240 blows/foot) at a depth of 

11 feet below excavation (27.5 feet below river bed). The nominal resistance at that 

depth is 579 kips. Maximum calculated driving stresses were 3.7 ksi discouraging 

the use of a larger hammer, unless a sufficiently high concrete strength would be 

chosen. If displacement piles were indeed used, predrilling or jetting would likely be 

required to advance the piles through the upper stratum.

A low displacement pile such as an H-pile or open end pipe pile, which would cause 

a lower or no densification, presents a more attractive foundation solution. Thus, the 

analysis was repeated for a 55 feet long, HP 14x102 H-pile which would allow the 

pile to reach bedrock at 53.5 feet depth (70 feet below river bed). Note that it is 

common practice to analyze H-pile drivability with a toe resistance as though the Hpile would be plugging while at the same time assuming a toe quake of only 0.10 

inches as for a non-displacement pile. In other words it is tacitly assumed that the 
172

pile is only partially plugging or that the plug slips during driving. This also means 

that there is no major densification effect when driving the H-pile.

The drivability results, shown in Figure 12-7, include the pile penetration resistance, 

maximum compression stresses and transferred energy for the normal (first pile) and 

later (densified) conditions. The concrete pile results are also shown in Figure 12-8. 

Along with the wave equation drivability results for the low displacement HP 14x102 

H-pile. Figure 12-8 allows for a comparison of penetration resistance and 

transferred energy and shows that the low displacement steel pile will drive with a 

significantly lower penetration resistance though the extremely dense sand and 

gravel layer.

Note that the penetration depth in these figures corresponds to the depth below 

cofferdam excavation level. The maximum penetration resistance calculated for the 

H-pile to penetrate the extremely dense sand and gravel stratum is only 45 blows 

blows/foot. Corresponding compression driving stresses do not exceed 32 ksi and 

are within driving stress limit. The nominal resistance increases significantly at 

deeper pile penetration depths. However, the penetration resistance increases to 

only 44 blows/foot before quickly transitioning to refusal conditions when the pile 

reaches rock. The results indicate that the H-pile could be driven to bedrock 

allowing for a possible higher factored load, reduced number of piles and a more 

cost effective design.
173

Figure 12-7 Nominal resistance, calculated blow count and stresses for before and 

after densification and nominal resistance for the after scour condition.

Figure 12-8 Drivability results for H-pile with corresponding concrete pile results.
174

12.5.4 Example 4 – Tension and Compression Stress Control

Example 4 illustrates the use of the wave equation for the control of tension stresses 

in a 60 feet long concrete pile. Static calculations indicate a 14-inch square, 

prestressed concrete pile driven through 20 feet of loose silty fine sand, 35 feet of 

medium dense sandy silt, and 3 feet into a dense sand and gravel deposit could 

develop a required nominal resistance of 400 kips. The static analysis also indicates 

that the nominal resistance is distributed as 55% shaft resistance and 45% toe 

resistance with a variable shaft resistance distribution along the pile shaft as shown 

in Figure 12-9. 

The contractor selected a Junttan HHK 3A hydraulic hammer for driving the 

prestressed concrete piles. This hammer has a ram weight of 6.6 kips and a rated 

energy of 26 ft-kips. The contractor’s hammer submittal indicates that the hammer 

cushion will consist of 8 inches of a material that has an elastic modulus of 360 ksi 

and a coefficient of restitution of 0.90 and a cross sectional area of 250 in2

. A 

helmet weighing 1.0 kip is also planned for the driving system. 

Figure 12-9 Example 4 – Problem profile. 
175

The concrete pile will have a compression strength of 5.5 ksi and an effective 

prestress after losses of 0.70 ksi. Using the AASHTO (2014) driving stress 

recommendations, discussed in Chapter 8, results in maximum recommended 

driving stresses of 4.1 ksi in compression and 0.92 ksi in tension.

One of the main concerns with the drivability of concrete piles is the possibility of 

developing high tension stresses during easy driving conditions when the soil 

provides little or no toe resistance. Therefore, the wave equation should be used to 

evaluate the contractor’s proposed driving system during both low and high 

resistance conditions. 

First, an evaluation of tension stresses during easy driving is presented. The weight 

of the pile and driving system is anticipated to be on the order of 20 kips. Hence, the 

pile penetration depth for the wave equation analysis should be selected below the 

depth to which the pile will likely penetrate or “run” under the weight of the pile and 

driving system, or approximately 10 feet. At this depth, the pile is still within the 

loose silty fine sand stratum and tension driving stresses are anticipated to be near 

their peak. Although not strictly correct, for the first low resistance analysis of 20 

kips it is accurate enough to assume the same shaft resistance percentage (55%) as 

for the final penetration of 58 feet. For a complete bearing graph not only the 20-kip

nominal resistance, but also other higher values are input and analyzed.

The contractor submitted a plywood pile cushion design comprised of four 3/4 inch 

sheets with a total thickness of 3 inches. Pile cushion stiffness significantly affects 

tension driving stresses. Therefore, it is necessary to determine whether or not the 

contractor’s proposed pile cushion thickness is sufficient to maintain tension stress 

levels below specified limits. In the first trial, the 3 inch pile cushion is assumed to 

possess the properties of new plywood. Thus, the original pile cushion thickness of 

3 inches and the new cushion elastic modulus of 30 ksi are input. Based on this 

information, the wave equation analysis indicates for the 20-kip nominal resistance a 

maximum tension stress of 1.47 ksi. The magnitude of the calculated tension stress 

exceeds the allowable driving stress limitation of 0.92 ksi.

A second wave equation analysis was therefore performed with an increased pile 

cushion thickness of 6 inches. By using the thicker pile cushion, the maximum 

tension stress was reduced from 1.47 ksi to 0.89 ksi which was less than the 

specified driving stress limit. The original and second wave equation analysis 

results for the easy driving condition at a pile penetration depth of 10 feet are 

presented in Figure 12-10. 
176

Figure 12-10 Bearing graphs for easy driving condition, two pile cushion 

thicknesses. 
177

The Junttan HHK 3A hydraulic hammer has an adjustable stroke height (or energy 

level). In the previous two analyses, it was assumed that the hammer would be run 

at the maximum rated energy of 26 ft-kips which corresponds to an equivalent stroke 

of 3.94 feet. To reduce pile cushion costs, the contractor suggested piles be driven 

with a reduced stroke and thinner less costly pile cushion. The wave equation 

analysis was therefore repeated for a 3 inch cushion and a 1.75 foot hammer stroke, 

corresponding to a potential energy of 11.6 ft-kips. For this configuration, the 

calculated maximum tension stress was 0.90 ksi at the 20-kip nominal resistance.

If the hammer would only be operated at the reduced 1.75 foot stroke, then the blow 

count would be at refusal (greater than 240 blow/foot) for the required 400 kip 

nominal resistance. Therefore, an additional analysis was performed for the final 

penetration depth of 58 feet using the full hammer stroke and rated energy and a 

used 3 inch thick pile cushion. The cushion was modeled with an elastic modulus of 

75 ksi which considers that the cushion thickness is also reduced. Obviously, in this 

case it is not necessary to analyze the 20 kip nominal resistance value.

The wave equation analysis results for the final driving condition are shown in Figure 

12-11. They indicate a penetration resistance of 56 blows/foot for the 400 kip 

required nominal resistance. The associated tension and compression stresses are 

0.58 and 3.32 ksi which are less than the maximum recommended driving stresses 

of 0.92 and 4.15 ksi, respectively.

Figure 12-11 also indicates that the tension stresses for the full stroke driving would 

be excessive, greater than 0.92 ksi, if full stroke driving were used with the 3-inch 

cushion at a nominal resistance of 250 kips or less. For this resistance value, on the 

other hand, the reduced stroke analysis indicates a penetration resistance of 85 

blows/foot. Based on these findings the following recommendation would be made to 

the contractor:

• Use a fresh 3-inch thick plywood cushion for every pile.

• Operate the hammer at a reduced equivalent stroke of 1.75 ft.

• Once the penetration resistance reaches 85 blow/foot, increase the 

equivalent stroke to the full rated stroke of 3.94 feet.

• Drive the pile to a minimum driving resistance of 56 blows/foot at the full rated 

stroke condition.

Note that it may be quicker, simpler, and possibly more cost effective to use the 6 

inch pile cushion and full hammer energy for the complete driving sequence.
178

Figure 12-11 Bearing graphs for early low energy and end of driving high energy 

driving condition. 
179

12.5.5 Example 5 – Use of Soil Setup

Consider the soil profile in Figure 12-12. In this example, a 12 inch square, 

prestressed concrete pile is to be driven into a thick deposit of stiff clay. The stiff 

clay has an average shear strength of 1.5 ksf. Based on field vane shear tests, it is 

estimated that the remolded shear strength at the time of driving will be 1.1 ksf, 

resulting in an expected soil setup factor of 1.36. A static analysis indicates that a 

nominal resistance of 300 kips after setup can be obtained for the proposed pile type 

at a penetration depth of 50 feet. The static analysis also indicates that the nominal 

resistance is distributed as 92% uniform shaft resistance and 8% toe resistance.

The contractor selected a Vulcan 08 single acting air hammer for driving the 

prestressed concrete piles. The contractor's hammer submittal indicates that the 

hammer cushion will consist of 8.5 inch of Hammortex with a cross sectional area of 

148 in2. The pile cushion will consist of plywood with a total thickness of 6 inches. It 

is anticipated that the pile cushion will compress and stiffen during driving similar to 

that described in Example 4. For an easy driving analysis, the assumption of a new 

pile cushion with the elastic modulus of 30 ksi would apply. For the late driving 

scenario, a compressed cushion modulus of 75 ksi should be considered. As 

mentioned earlier, using this 2.5 times higher modulus will not only account for the 

stiffening of the material but also for its thickness decrease. The contractor's 

submittal indicates that the helmet weighs 2.6 kips.

Based upon the reported soil type and setup behavior, a 36% increase in nominal 

resistance with time is expected at this site. Therefore, piles could be driven to a 

reduced nominal resistance, or static resistance to driving (SRD), of 225 kips instead 

of the required value of 300 kips with the remaining 75 kips expected from soil setup. 

As noted earlier in this chapter, a static load test will be performed on the project to 

confirm the expected resistance gain.
180

Figure 12-12 Example 5 –- Problem profile. 

The wave equation results presented in Figure 12-13 indicate a final penetration 

resistance of 43 blows/foot could be used as the driving criteria for a 225 kip SRD or 

end of driving nominal resistance. This is significantly less than the 89 blows/foot 

required for a nominal resistance of 300 kips. Hence, significant pile length and 

driving effort may be saved by driving the piles to the lower resistance. However, 

this approach requires a restrike test or a static load test sometime after pile 

installation; the waiting time period is soil type dependent as discussed in Section 

7.2.4.
181

Figure 12-13 Using a bearing graph with soil setup. 
182

12.5.6 Example 6 – Driving System Characteristics 

Example 6 presents a wave equation comparison of two hammers having the same 

potential energy. Dynamic formulas such as the Modified Gates formula consider 

only the potential energy of the driving system. Therefore, the penetration 

resistance required for a specific nominal resistance by this and most other dynamic

formulas would be the same for two hammers provided that the hammers had the 

same potential energy. In this example, the penetration resistances predicted by the 

wave equation for these two hammers in the same pile-soil condition is, however, 

quite different.

In this example, a 14 inch O.D. x 0.375 inch wall, closed end pipe pile is to be driven 

to a nominal resistance of 400 kips. The pile has a furnished length of 66 feet and 

an embedded length of 52.5 feet. A static analysis indicates that the soil resistance 

distribution will be 30% shaft resistance and 70% toe resistance. The shaft 

resistance will be distributed triangularly along the embedded portion of the pile 

shaft. The example problem's soil profile is presented in Figure 12-14. With a very 

dense, dry soil at the pile toe, the normal program recommendation for the quake at 

the pile toe is D/120 or 0.12 inch. However, experience with the bearing layer from 

previous dynamic measurements showed that the silty fine sand at this site is highly 

elastic and has a larger than normal toe quake of 0.40 inch.

The contractor is considering using either a Vulcan 014 air hammer or an ICE 42-S 

open end diesel hammer to drive the piles. Both hammers have the same 

manufacturer’s rated hammer energy of 42 ft-kips. However, the ram of the Vulcan 

014 is roughly 3.5 times heavier than the ram of the ICE hammer. Details of these 

hammers and their associated proposed driving systems are summarized in Table 

12-2. 

Wave equation results for the two hammers are plotted on the same bearing graph 

in Figure 12-15. For the high toe quake case (0.40 inches), wave equation analysis 

calculates a penetration resistance of 94 blows/ft to achieve a 400 kip nominal 

resistance with the heavy ram air hammer whereas the lighter ram diesel hammer

requires a penetration resistance of 209 blows/ft. For the standard toe quake case 

(0.12 inches), the heavy ram air hammer requires a penetration resistance of 56 

blows/ft while the light ram diesel hammer requires 99 blows/ft. Hence, even though 

both hammers have the same potential energy, the required penetration resistance 

for the 400 kip nominal resistance is quite different.
183

Figure 12-14 Example 6 – Problem profile.

Table 12-2 Example 6: Proposed Hammer and Driving Systems

Hammer Model Vulcan 014 ICE 42S

Ram Weight, kips 14 4.09

Rated Energy, ft-kips 42 42

Rated Stroke, ft 3.0 10.3

Helmet Weight, kips 1.67 2.05

H. Cushion Material Nycast Blue Nylon

H. Cushion E-Mod, ksi 208 175

H. Cushion Area, in2 234 398

H. Cushion Thickness, in 6.0 2.0

H. Cushion COR 0.91 0.92
184

Figure 12-15 Bearing graph – for two hammers with equivalent 

potential energy and large toe quake.
185

Even though the Vulcan 014 requires a lower penetration resistance (blow count) for 

the same nominal resistance and has a lower efficiency (0.67 vs. 0.80 for the diesel 

hammer), it transfers roughly 20% more energy into the pile. This is because, first, 

the diesel hammer uses part of its energy to compress the gasses prior to impact. 

Second, the lower impact velocity of the heavy hammer is associated with lower 

energy losses. And lastly, the duration of the air hammer’s impact is longer and 

consequently more effective at driving into a highly elastic soil with a large quake.

It is, however, interesting to note that for the smaller normal quake case in Figure 

12-16, the lighter ram’s blow counts improve relative to the heavier hammer at high 

nominal resistances and blow counts. This phenomenon can be explained with the 

diesel hammer’s higher stroke and, therefore, higher impact force during harder 

driving. At the higher penetration resistance levels, energy is not as important as 

force to overcome the soil resistance.

This example illustrates the dynamic complexities of hammer-pile-soil interaction. 

Clearly, the potential energy alone, which is the sole hammer input in dynamic 

formulas, does not adequately assess pile drivability.
186

Figure 12-16 Bearing graph – for two hammers with equivalent 

potential energy and normal toe quake.
187

12.5.7 Example 7 – Assessment of Pile Damage 

Another pile driving construction concern is pile damage. Although it is frequently 

assumed that steel H-piles can be driven through boulders and fill materials 

containing numerous obstructions, pile installation reviews reveal that this

assumption is invalid. H-piles without commercially manufactured pile toe 

reinforcement present one of the most commonly damaged pile types. The damage 

occurs because of the ease with which flanges can be curled, rolled, and torn. 

Because deforming a pile plastically or otherwise non-elastically requires a 

significant amount of energy, pile damage has a detrimental effects on both 

penetration resistance and, therefore, nominal resistance activation (the blow count 

will be higher at the same resistance).

Figure 12-17 Example 7 –- Problem profile.

This example illustrates how the wave equation can be used to obtain insight into a 

driving situation involving pile damage. The project conditions are shown in Figure 

12-17. The HP 12x53 H-piles are 40 feet in length with a factored load of 304 kips 

and a nominal resistance of 380 kips based on confirmation with a static load test. 
188

The soil profile consisted of 15 to 17 feet of miscellaneous fill, including some bricks 

and concrete. Below the fill, 15 feet of silty clay overlay bedrock which was 

encountered at a depth of about 35 feet. The design called for driving the piles to 

bedrock.

The contractor selected a Pileco D12-42 single acting diesel hammer with a rated 

energy of 29.9 ft-kips to drive the piles. Using the FHWA modified Gates formula 

specified in the contract documents, the required penetration resistance was 46

blows/ft with this hammer for a nominal resistance of 380 kips.

Figure 12-18 presents the wave equation results indicating a nominal resistance of 

245 kips at the Gates blow count of 46 blows/ft, well below the required 380 kip 

nominal resistance. On the other hand, the wave equation also showed that the 

maximum compression stresses at the pile toe would reach 40 ksi when the required 

nominal resistance was reached at 104 blows/foot and even 43 ksi at refusal. Most 

H-pile sections are now made of steel with a 50 ksi yield strength. However, in this 

example case, the yield strength was only 36 ksi and the compression driving stress

limit only 32.4 ksi (90% of the yield strength).

In accordance with the contract requirement, several static load tests were 

conducted. In all cases, the piles failed to carry the 380 kip nominal resistance in 

spite of the fact that several of the piles were eventually driven to a penetration 

resistance exceeding 240 blows/ft with no indication of damage at the pile head. 

Because of the high penetration resistances to which several piles were driven, it 

was apparent that even harder driving would not result in a higher nominal 

resistance. Consequently, the contractor was requested to pull several of the piles 

to check for possible damage. Upon extraction, it was noted that the piles were 

severely damaged at the pile toe. The flanges were separated and rolled up from 

the web. While the damage probably occurred as the unprotected piles were driven 

through the miscellaneous rubble fill, it is also obvious from Figure 12-18 that the 

refusal blow count would generate dynamic steel pile stresses in excess of 40 ksi, 

and therefore in excess of the driving stress limit. The highest stresses would occur 

at the pile toe according to the numerical wave equation results while the stresses at 

the pile top were significantly lower (thus no damage at the pile head). Tables 12-3

and 12-4 show the maximum values of forces, stresses, and other variables 

calculated by wave equation analysis for both the required nominal resistance of 380 

kips and the refusal driving situation (penetration at or in excess of 240 blows/ft) 

respectively.
189

Figure 12-18 Wave equation bearing graph for proposed driving system.
190

Table 12-3 Example 7 – Maxima of Forces, Stresses and Other Variables for 

Required Nominal Resistance of 380 kips

Segment 

No.

Max 

Tension 

Force 

kips 

Max 

Compr. 

Force 

kips 

Max 

Tension 

Stress 

ksi 

Max 

Compr. 

Stress 

ksi 

Max 

Velocity 

ft/s 

Max 

Displacement 

inches

Max 

Transfer 

Energy 

kip-ft 

Top = 1 0.0 502.2 0.00 32.40 16.49 0.584 15.99

2 -18.9 492.1 -1.22 31.75 16.60 0.556 15.61

3 -35.7 498.7 -2.31 32.17 16.56 0.524 15.07

4 -49.1 501.2 -3.17 32.34 16.49 0.490 14.35

5 -54.0 502.4 -3.48 32.41 16.29 0.454 13.46

6 -53.9 499.6 -3.48 32.23 16.06 0.415 12.42

7 -50.7 495.6 -3.27 31.98 15.77 0.374 11.23

8 -39.6 502.7 -2.56 32.43 15.43 0.335 10.13

9 -31.3 523.0 -2.02 33.74 15.17 0.301 9.24 

10 -27.6 537.3 -1.78 34.66 15.02 0.272 8.57

11 -21.7 556.4 -1.40 35.90 14.01 0.244 7.96

Toe = 12 -12.2 613.9 -0.79 39.61 9.58 0.216 7.46

Table 12-4 Example 7 – Maxima of Forces, Stresses and Other Variables for at 

Refusal Driving Condition (Nominal Resistance of 500 kips) 

Segment 

No.

Max 

Tension 

Force 

kips 

Max 

Compr. 

Force 

kips 

Max 

Tension 

Stress 

ksi 

Max 

Compr. 

Stress 

ksi 

Max 

Velocity 

ft/s 

Max 

Displacement 

inches

Max 

Transfer 

Energy 

kip-ft 

Top = 1 0.0 586 0.00 37.80 17.51 0.590 17.39

2 -23.6 539.2 -1.52 34.79 17.63 0.559 16.94

3 -40.7 530.5 -2.63 34.23 17.61 0.526 16.32

4 -59.3 532.2 -3.82 34.33 17.46 0.490 15.48

5 -69.1 532.5 -4.46 34.36 17.22 0.452 14.43

6 -69.6 540.2 -4.49 34.85 16.90 0.411 13.19

7 -57.0 552.2 -3.68 35.62 16.50 0.367 11.77

8 -53.0 566.0 -3.42 36.52 16.09 0.320 10.23

9 -50.2 582.3 -3.24 37.57 15.69 0.271 8.65

10 -40.8 591.6 -2.63 38.17 15.33 0.224 7.27

11 -23.6 613.5 -1.52 39.58 13.79 0.184 6.23

Toe = 12 -13.3 661.6 -0.86 42.68 8.57 0.147 5.48
191

The effect of the damage on the pile drivability can also be evaluated with a wave 

equation analysis. Since static load tests indicate that piles driven as hard as 240 

blows/foot did not support the 380-kip load, one pair of nominal resistance and 

penetration resistance values is available as a reference point on the wave equation 

bearing graph. For the damaged pile scenario, a bearing graph may be calculated 

for the pile with damaged toe by a simply reducing the elastic modulus of the lowest 

pile segment until results agree with the penetration resistance and nominal 

resistance observations. In the present case, the reduced toe segment stiffness was 

found to be approximately 10% of the normal value.

Figure 12-19 presents wave equation results for both the undamaged and the 

damaged pile toe scenarios. The results indicate that the nominal resistance of 380 

kips could not be obtained for the damaged pile, regardless of the penetration 

resistance. Essentially, the damaged pile section "cushioned" the hammer blow at 

the pile toe and attenuated the hammer energy. Once damaged, the soil resistance 

at the pile toe could not be overcome, and therefore, the pile toe would not advance. 

The above illustrates that not only blow counts but also driving stresses also may 

limit the drivability of a pile to the required nominal resistance.

The potential for pile damage on this project could have been greatly reduced if a 

wave equation had been performed during the design stage or had been specified 

for construction control. As pointed out earlier, the wave equation bearing graph in 

Figure 12-19 illustrates that the nominal resistance of 380 kips could only be 

obtained by the contractor's driving system at a penetration resistance of 104 

blows/foot or more with an associated pile toe stress of 40 ksi, a stress in excess of 

the steel yield strength of 36 ksi and recommended limit driving stress of 32.4 ksi.

Considering that the stresses calculated by the wave equation are averages over the 

cross section, a non-uniform distribution of the soil or rock resistance could have 

added significant additional bending stresses in the steel pile near its toe. Hence, 

the potential damage would have been clearly apparent at the time of the 

contractor's hammer submittal had a wave equation analysis been performed. 

Additional wave equation analyses of the contractor's driving system could have 

been performed at the same time to determine if driving stress levels could be 

acceptably reduced by using reduced fuel settings or shorter hammer strokes. If 

driving stresses could not be controlled in this manner, as is most likely in the 

present case, approval of the proposed driving system should not have been 

obtained, and either alternate hammers should have been evaluated or a higher 

steel yield strength required.
192

Figure 12-19 Wave equation bearing graphs for damaged and undamaged pile.
193

In any event, where H-piles have to be driven through materials that could include 

obstructions or where piles have to be driven to hard rock, it is always strongly 

recommended to protect the pile toe with a so-called driving shoe. Driving shoes are 

discussed in Chapter 16 and consist of steel castings which can be welded to the 

pile toe. Driving shoes tend to centralize the toe resistance force and/or reinforce 

the flanges.

12.5.8 Example 8 – Selection of Wall Thickness

This wave equation example demonstrates the selection process for the required 

wall thickness of a pipe pile. Consider the soil and problem profile presented in 

Figure 12-20. The foundation details are based on a bridge construction situation 

where a major bridge was built over a lake with a water depth of 85 feet. The piles 

extended 25 feet above the water level. Based upon static soils analysis the piles 

were to be driven through 25 feet of granular overburden, 5 feet of overconsolidated 

glacial till, and then to a hard bedrock. Thus total pile length was 140 feet. The piles 

had to be designed for a factored load of 1600 kips; a resistance factor of 0.80 

(static load testing was required) yielded a required nominal resistance of 2000 kips. 

Structural considerations called for a minimum wall thickness of 0.50 inches. 

However, as a drivability check, wave equation analyses were performed for the 48 

inch outside diameter pipe pile with four different wall thicknesses of 0.50, 0.75, 

1.00, and 1.25 inches. The analyses were based on the following assumptions:

1. A Berminghammer 6005 open end diesel hammer would be available. This 

hammer has a ram weight of 13.6 kips and a rated energy of 161 ft-kips. 

2. The pile is driven open ended and because of its relatively shallow 

penetration, it will not plug.

3. The toe resistance will develop against only the steel toe area on the hard 

rock. The toe resistance is estimated to be 1600 kips or 80% of nominal 

resistance for all four pile types regardless of wall thickness.

4. Because the pile is a low-displacement type driven to hard rock, the program 

recommended toe quake is 0.04 inches. (For moderately hard rock the 

standard 0.10 inch quake could be assumed.)

5. Average shaft damping was assumed to be 0.10 s/ft.
194

6. During driving, open end pipe piles with diameters greater than 30 inches 

usually do not plug and for moderate penetrations (say less than 10 

diameters) it may be assumed for the dynamic analysis that at least partial 

internal friction occurs. The shaft resistance in the present case was 

assumed to act equally on the inside and outside surfaces of the pile and 

amount to 20% of the total nominal resistance. For greater pile penetration 

depths, it would be less likely that so much internal friction occurs during 

driving because the effective stresses acting inside the pipe and, therefore, 

the unit shaft resistance inside the pipe would be less than for the outside 

shaft resistance. For that reason, the internal friction is often only assumed to 

be either 0 or 50% of the outside friction. In wave equation analysis 

modeling, a 50% inside friction could be simulated by a 50% increase in the 

pile perimeter while the unit shaft resistance would remain as calculated for 

the outside pile wall.

Figure 12-20 Example 8 – Problem profile.
195

Figure 12-21 presents the results of these analyses in the form of bearing graphs 

and Table 12-5 summarizes calculated stresses, blow counts, stroke and transferred 

energy for the required 2000 kip nominal resistance.

Calculated maximum driving stresses are not more than 43 ksi for the 0.50 inch wall 

pipe and less than 35 ksi for the other thicker wall pile sections. Assuming that the 

pipe piles would be manufactured from Grade 3 pipe with yield strength of 45 ksi, the 

driving stress limit would be 40.5 ksi. Therefore, all except the 0.5 inch wall piles 

should withstand the driving stresses. However, uneven or sloping rock would 

cause higher localized stress concentrations. It would, therefore, be wise to choose 

a greater wall thickness for which the driving stresses allow for an additional margin 

of safety.

Table 12-5 Example 8 – Wave Equation Results for Four Potential Pipe Pile Wall 

Thicknesses at 2000 kip Nominal Resistance 

Wall 

Thickness 

inches

Stroke 

feet

Transferred 

Energy

ft-kips 

Compression 

Stress 

ksi 

Blow Count 

blows/ft 

0.50 9.1 93 42.8 127

0.75 8.7 79 34.2 100

1.00 8.5 71 29.1 90

1.25 8.6 68 25.4 88

For hammer approval, AASHTO (2010) LRFD Bridge Construction Specification 

require the wave equation penetration resistance to be between 24 and 120 

blows/foot. The wave equation analysis results indicate that the 0.50 inch thick wall 

pipe would require too high a blow count while the 0.75 inch wall thickness would 

have an acceptable penetration resistance. Use of the 1.00 or 1.25 inch wall 

thickness would help little with either drivability or driving stresses and be much 

more costly. Table 12-5 also shows that while hammer strokes were nearly the 

same, the transferred energies varied significantly with the more flexible piles 

accepting more energy. However, this energy was needed to elastically compress 

the very long piles and was therefore not available for actual work to overcome the 

soil resistance.
196

Figure 12-21 Bearing graphs for open end pipe piles with wall thicknesses

of 0.50, 0.75, 1.00, and 1.25 inches.
197

The actual project which was used as a guide in setting up this example was 

successfully completed with the 0.75 inch wall thickness. Reportedly, the static load 

test did not fail under a 2000-kip proof test. While this project was successfully 

completed, a caution about driving and analyzing large diameter open end pipe piles

should be added. When driving to dense soil layers, large pipes may not plug during 

driving and they may plug under static loading conditions. Pile driving may therefore 

be much easier than predicted by the wave equation if the analysis were made on 

the assumption of a plugged pile which behaves like a displacement pile with high 

toe resistance and a large toe quake. The situation may also be different at the end 

of driving and during a restrike. It is recommended to perform wave equation 

analyses with upper and lower bound soil resistance values to assess hammer 

sizing.

It is also important to remember that a driving criterion for a pile driven to hard rock 

should not be specified in terms of blows per unit penetration, but rather as a 

maximum penetration for a certain number of hammer blows. In the present case, 

for the pile with the 3⁄4 inch wall (a required blow count of 100 blows per foot was 

calculated for the 3⁄4 inch wall pile) it would be reasonable to require that the pile be 

accepted when its penetration under 10 consecutive hammer blows is less than 0.5 

inches while the hammer stroke is between 8.5 and 9.0 ft. Specifying that the pile 

would have to penetrate a certain distance into a hard rock to qualify for having 

achieved the nominal resistance would most probably lead to pile toe damage. 

Even when driving into moderately hard rock, overdriving the piles should be done 

cautiously. This is sometimes problematic, since some shales or weathered rock 

materials exhibit relaxation after driving and it is desirable to drive piles in those 

formations to a capacity in excess of the required nominal resistance. Wave 

equation results should indicate that the higher resistance can be achieved within 

material stress limits if overdriving is required.

12.5.9 Example 9 – Evaluation of Vibratory Driving

This example illustrates the use of a wave equation analysis for evaluating vibratory 

hammer installation of the sheet piles required for the cofferdam construction. The 

situation is depicted in Figure 12-22 and is similar to Example 3. The sheet piles 

must be installed using a vibratory hammer. The contractor has an ICE 815 hammer 

available and intends to drive pairs of AZ18 sheet piles whose combined cross 

sectional area is 29.5 in2. These are Z-section sheets, each with a width of 24.8 

inches, a depth of 15 inches and a thickness of 0.37 inches. At the time of sheet pile 

installation, the soil within the cofferdam is not excavated. The 50 foot long sheet 

piles are vibrated from mudline to an estimated depth of 35 feet below mudline.
198

Figure 12-22 Example 9 – Problem profile.

For the non-excavated condition, the sheet piles must first penetrate a 16.5 foot thick 

layer of soft silt, followed by very dense sand and gravel, and then dense sand and 

gravel layers. The static resistance values were calculated based on corrected SPT 

values of 5 for the silt, 110 for the very dense sand and gravel, and 33 for the dense 

sand and gravel. As is reasonable for submerged coarse grained soils subjected to 

vibratory driving, a shaft resistance loss of 75% of the long term resistance was 

assumed while no loss of resistance was assumed for the silt, which was considered 

cohesive. This situation is modeled with respective setup factors for silt and sand 

layers of 1.0 and 4.0. The associated gain/loss factor was, therefore, set to 0.25 

which produces an SRD in the sands which is 25% of the long term resistance while 

the silt is assumed having an SRD equal to the long term resistance. (Note: it is not 

recommended to use these setup and gain/loss factors for drivability analyses of 

impact pile driving). For the toe resistance it was conservatively assumed that the 

full long term resistance would be present during the vibratory driving (gain/loss 

factor equal to 1.0).

As per program recommendations, damping factors were set to twice the values 

assumed for impact driving, i.e., a shaft damping of 0.30 s/ft in the silt (assumed to 

behave cohesive) and 0.10 s/ft in the sand and gravel layers. Toe damping was 
199

input as 0.30 s/ft for all layers. Also the damping type was switched from Smith to 

Smith-viscous per program recommendations. In addition, program 

recommendations are that only quakes should be doubled for cohesive soils but not 

non-cohesive soils. This guidance was followed for the silt layer while quakes were 

left at their 0.10 inch default values in the sands and gravels. The calculated unit 

soil resistance values and the associated dynamic soil parameters are shown in 

Table 12-6. 

Table 12-6 Soil Information for Vibratory Sheet Pile Driving in Example 9

Depth 

feet

Unit Shaft 

Resist 

ksf 

Unit Toe 

Resist. 

ksf 

Skin 

Quake 

inch 

Toe 

Quake 

inch 

Skin 

Damping 

s/ft 

Toe 

Damping 

s/ft 

0.0 0.0 0.0 0.20 0.20 0.30 0.30 

16.5 0.283 22.7 0.20 0.20 0.30 0.30 

16.5 0.465 250.6 0.10 0.10 0.10 0.30 

29.0 1.272 250.6 0.10 0.10 0.10 0.30 

29.0 0.274 137.8 0.10 0.10 0.10 0.30 

43.0 0.453 137.8 0.10 0.10 0.10 0.30 

The pertinent information needed for the hammer model is shown in Figure 12-22, 

i.e., bias mass of 5.5 kips and an oscillator mass of 8.1 kips of the ICE 815 hammer. 

The clamp weight of 2.2 kips is added to the oscillator mass in the analysis. The two 

main masses are connected by elastomers which are modeled with very soft springs 

so as to isolate the crane line from the oscillator vibrations while at the same time 

adding crowd force to the driving system. The eccentric moment, represented in the 

program by a mass and a radius, is 4.4 inch-kips for the ICE 815 hammer. With a 

rated frequency of 26.7 Hz (1600 rpm) the centrifugal force of this unit is 320 kips. 

However, for conservatism, it is assumed that the hammer will only be run at 20 Hz 

which yields a centrifugal force of 180 kips. Efficiency and start-up time (the time 

necessary for the hammer to reach full frequency) are left at their respective 1.0 and 

0.0 default values. Another input is a 7-kip line pull, or upward directed crane force, 

which may be needed to maintain hammer-pile system stability. Oftentimes after the 

pile has sufficient embedment for stability, the operator will let the line slacken which 

will allow for a greater downward force and therefore an increase in the speed of pile 

penetration. An upward directed (positive) line force is therefore a conservative 

input.
200

Figure 12-23 shows the drivability results in graphic form and Table 12-7 shows the 

results in numerical form. For the first analyzed depth of 6 feet, the SRD is less than 

all of the applied weights (hammer weight plus clamp weight plus pile weight minus 

line pull) causing the sheet pile to “run” as indicated in the final results by the zero 

(0) penetration time. After the pile penetrates into the very dense sand layer, the 

required penetration time sharply increases to values up to almost 20 s/ft, but 

reduces to much more comfortable values as the sheet pile toe enters the dense 

sand and gravel. The final foot of penetration requires less than 5 s/ft. At that point, 

the SRD ranges between 85 and 90 kips with 28 kips acting at the sheet pile toe (the 

steel area of the pile). The calculated compression and tension stress maxima are 

rather small varying between 2 and 5 ksi. These stresses are typical for vibratory 

pile driving, by it should be remembered that these stresses are averaged over the 

pile cross sectional area. Local stress concentrations of much higher magnitudes 

must be expected, for example in the areas surrounding the pile clamp.

Vibratory hammer refusal has occasionally been specified as low as 1 inch/min 

corresponding to a very high penetration time of 720 s/ft, and the results, therefore, 

suggest that the sheet pile installation should be possible with the 815 hammer. 

However, the accuracy of the wave equation prediction strongly depends on the 

realism of the relatively crudely estimated static resistance to driving (SRD). 

Furthermore, a good alignment of the sheet piles and thus no excessive interlock 

friction is another condition for a successful installation.

It should be emphasized that the drivability analysis of vibratory pile driving is a 

reasonably reliable tool for equipment selection. On the other hand driving time 

estimates can widely fluctuate (even from pile to pile on the same) since they are 

extremely sensitive to soil resistance variations, particularly at the pile toe. Even 

greater uncertainty exists when attempting to relate SRD and even more so, long 

term resistance to penetration time. While research in this area has been and is 

occasionally being done, no really reliable methods have yet been established that 

would allow for a nominal resistance calculation from vibratory driving observations.
201

Figure 12-23 Nominal resistance and wave equation calculated penetration time for 

vibratory drivability analysis.
202

Table 12-7 Nominal Resistance and Wave Equation Calculated Results for 

Vibratory Drivability Analysis. Total Drive Time – 4 Minutes 

Depth 

feet

Nominal 

Resist. 

kips 

Shaft 

Resist. 

kips

Toe 

Resist. 

kips

Pen. 

Time 

s/ft 

Comp. 

Stress 

ksi 

Tension 

Stress 

ksi 

6.0 5.1 3.5 1.7 0.0 0.0 0.0

12.0 17.2 13.8 3.4 0.9 2.4 -1.6

13.0 19.9 16.3 3.6 1.0 2.5 -1.5

14.0 22.8 18.8 3.9 1.1 1.9 -1.4

15.5 27.4 23.1 4.3 1.2 1.9 -1.3

17.5 78.7 27.6 51.1 15.5 3.1 -3.4

19.0 81.1 30.0 51.1 15.7 3.2 -3.5

20.0 83.0 31.9 51.1 16.0 3.3 -3.6

21.0 85.0 33.9 51.1 16.4 3.4 -3.7

22.0 87.2 36.1 51.1 16.7 3.5 -3.8

23.0 89.6 38.5 51.1 16.9 3.6 -3.9

24.0 92.1 41.0 51.1 17.5 3.8 -4.0

25.0 94.9 43.8 51.1 17.9 4.0 -4.2

26.0 97.8 46.7 51.1 18.2 4.1 -4.3

27.0 100.9 49.8 51.1 18.9 4.3 -4.5

28.0 104.2 53.1 51.1 19.5 4.5 -4.7

30.0 85.5 57.4 28.1 4.5 4.3 -4.0

31.0 86.3 58.2 28.1 4.5 4.3 -4.1

32.0 87.2 59.1 28.1 4.5 4.4 -4.2

33.0 88.1 60.0 28.1 4.6 4.4 -4.2

34.0 89.0 60.9 28.1 4.6 4.5 -4.3

35.0 90.0 61.9 28.1 4.6 4.6 -4.4
203

12.6 ANALYSIS DECISIONS FOR WAVE EQUATION MODELING

A wave equation analysis offers the engineer a variety of input and analysis options. 

Choosing just one of these options may limit what can be learned from the 

simulations. For example, it may be helpful to calculate both a bearing graph and 

drivability plot or investigate pile stress extrema and blow counts for both a diesel 

hammer and a hydraulic one. The two hammer types would be different as far as 

energy transfer and driving resistance even if the hammers are identically rated. 

The drivability analysis would show certain potential driving problems such as startup problems for the diesel in easy driving or refusal in dense intermediate soil layers. 

The following sections aim at helping the analyst to understand why certain options 

were made available and how they can help achieve an optimal pile installation.

12.6.1 Selecting the Proper Approach

Even though the wave equation analysis is an invaluable tool for the pile design 

process, it should not be confused with a static geotechnical analysis. Some wave 

equation programs, provide a simplified static analysis for resistance distribution 

purposes. However, the basic wave equation approach does not determine the 

nominal resistance of a pile based on soil boring data. The wave equation calculates 

a penetration resistance for an assumed nominal resistance (ultimate capacity), or 

conversely, it assigns an estimated nominal resistance (ultimate capacity) to a pile 

based on a field observed penetration resistance. It is one thing to perform a wave 

equation bearing graph for an expected nominal resistance at a particular pile 

penetration and a totally different matter to actually realize that nominal resistance at 

that depth. Significant cost overruns can occur when pile lengths required during 

construction vary significantly from those estimated during design by a static 

analysis. To avoid such occurrences, it is imperative that a static analysis, as 

described in Chapter 7, precede the wave equation analysis. The static analysis will 

yield an approximate pile penetration for a desired nominal resistance or a nominal 

resistance for a certain depth. The static analysis can also generate a plot of 

estimated pile nominal resistance as a function of depth. As a preparation for the 

wave equation analysis, it is important that the static analysis evaluates the soil 

resistance in the driving situation (e.g., remolded soil strength, before excavation, 

before scour, before fill placement, etc.). For the assessment of long term static 

conditions, the static analysis must consider the influence of soil setup or relaxation, 

additional change due to excavation, water table variations, and scour, etc.

After completion of the static analysis, a wave equation analysis should be 

performed, leading to either a bearing graph as illustrated in Example 1, or a 
204

drivability analysis of penetration resistances and stresses versus depth as 

presented in Example 3. Sometimes both analyses are performed. The validity of 

the bearing graph depends on the proximity of the analyzed soil profile and the site 

variability of the soil properties. The drivability analysis calculates penetration 

resistances and stresses for a number of penetration depths and, therefore, provides 

a more complete result. However, there is a very basic difference between these 

two approaches. The bearing graph approach allows the engineer to assess the 

nominal resistance of a pile given a penetration resistance at a certain depth and to 

formulate a driving criterion for a required nominal resistance. The drivability 

analysis points out certain problems that might occur during driving prior to reaching 

the target penetration. If the pile actually drives differently from the wave equation 

predictions, a reanalysis with different soil resistance parameters would be needed 

to match the observed behavior.

Even if an accurate static analysis and a wave equation analysis have been 

performed with realistic soil parameters, the experienced foundation engineer would 

not be surprised if the penetration resistance during pile installation were to differ 

substantially from the predicted one. Most likely, the observed penetration 

resistance would be lower than calculated. As an example, suppose that a pile had 

to be driven into a clay for a factored load of 182 kips. With a resistance factor of 

0.65 (dynamic testing of at least 2% of the piles would be specified), the required 

nominal resistance would be 280 kips. The static soil analysis indicates that the pile 

should penetrate to a depth of 82 feet to meet this nominal resistance requirement. 

There would be negligible toe resistance, and based upon remolded soil strength 

parameters, the soil may exhibit only 50% of its long term strength during driving 

(soil setup factor = 2). It is therefore only necessary to drive the pile to a nominal 

resistance of 140 kips, which should be achieved at the 82 foot depth. The expected 

end of installation penetration resistance would then correspond to 140 kips. A 

restrike test, performed 7 days after installation, would include soil setup effects and 

might show the required 280-kip nominal resistance and a much higher penetration 

resistance than at the end of driving.

The above discussion points out one major reason for differences between analysis 

and reality. However, as with all mathematical simulations of complex situations, 

agreement of wave equation results with actual pile performance depends on the 

realism of the method itself and on the accuracy of the model parameters. The 

accuracy of the wave equation analysis will be poor when either soil model or soil

parameters inaccurately reflect the actual soil behavior and when the driving system 

parameters do not represent the state of maintenance of hammer or cushions. The 
205

pile behavior is satisfactorily represented by the wave equation approach in the 

majority of cases. A review of potential wave equation error sources follows.

12.6.2 External Combustion Hammer Consideration

For external combustion hammers (e.g., air or hydraulically powered hammers) one 

of the most important wave equation input quantity is the hammer efficiency. It is 

defined as that portion of the potential ram energy that is available in the form of 

kinetic ram energy immediately proceeding the time of impact. Many sources of 

energy loss are usually lumped into this one number. If the hammer efficiency is set 

too high, an optimistically low penetration resistance would be predicted. This in 

turn could lead to a dangerous overprediction of nominal resistance. If the efficiency 

is set very low, for conservative pile nominal resistance assessments, the stresses 

may be underpredicted, leading to possible pile damage during installation.

Hammer efficiency and hammer stroke should be reduced for inclined (battered) pile 

driving. These reductions depend on the hammer type and batter angle. For 

hammers with internal ram energy measurements, no reductions are required to 

cover losses due to inclined pile driving. Modern hydraulic hammers often allow for 

a continuously adjustable ram kinetic energy which is measured and displayed on 

the control panel. In this case, the hammer efficiency need not cover friction losses 

of the descending ram but only losses that occur during the impact (e.g., due to 

improper ram-pile alignment), and it may therefore be relatively high (say 0.95). For 

such hammers, the wave equation analysis can select the proper energy level for 

control of driving stresses and economical penetration resistances by trying various 

energy (equivalent stroke) values that are lower than the rated value.

Similarly, a number of air/steam hammers can be fitted with equipment that allows 

for variable strokes. The wave equation analysis can help to find the penetration 

resistance at which the stroke can be safely increased to maximum. It is important, 

however, to realize that in fact the reduced stroke is often exceeded and the 

maximum stroke not fully reached. Corresponding increases and decreases of 

efficiency to cover the uncertainty of the actual equivalent stroke, may, therefore, be 

investigated.
206

12.6.3 Diesel Hammer Considerations

The diesel hammer stroke increases when the soil resistance, and therefore 

penetration resistance, increases. Certain wave equation programs simulate this 

behavior by trying a down stroke and, when the calculated up stroke is different, 

repeating the analysis with the new value for the down stroke until the strokes 

converge. The accuracy of the resulting stroke is therefore dependent on the 

realism of the complete hammer-pile-soil model and should be checked in the field 

by comparison with the actual stroke. The consequences of an inaccurate stroke 

could be varied. For example, an optimistic assumption of combustion pressure 

could lead to high stroke predictions and, therefore, non-conservative predictions of 

nominal resistance while stress estimates would be conservatively high (which may 

lead to a hammer rejection).

Stroke and energy transferred into the pile appear to be closely related, and large 

differences (say more than 10%) between stroke predictions and observations 

should be explained. Unfortunately, higher strokes do not always mean higher 

transferred energy values. When a diesel hammer preignites, probably because of 

poor maintenance, the gases combusting before impact slow the speed of the 

descending ram and cushion its impact. As a result, only a small part of the ram 

energy is transferred to the pile. A larger part of the ram energy remains in the 

hammer producing a high stroke. If, in this case, the combustion pressure would be 

calculated by matching the computed with the observed stroke under the 

assumption of a normally performing hammer, the calculated transferred energy 

would be much higher than the measured one, and the calculated penetration 

resistances (blow counts) would be non-conservatively low. It is, therefore, 

recommended that hammer problems are corrected as soon as detected on the 

construction site. If this is not possible, several diesel stroke or pressure options 

should be tried when matching wave equation results with field observations and the 

most conservative results should be selected. Section 12.8 discusses the available 

diesel hammer stroke options in greater detail.

Generally the hammer data file of wave equation programs contain reduced 

combustion pressures for those hammers which have stepwise adjustable fuel 

pumps. Note that decreasing combustion pressures may be associated with 

program input fuel pump settings that also have decreasing numbers. (For example, 

Delmag, APE D-Series, Pileco, ICE I-Series and other makes have fuel pump 

settings with 4 (maximum), 3, 2, and 1 (minimum), roughly correspond to 

combustion pressures of 100, 90, 81 and 73 percent of that associated with the 

hammer’s rated energy.) Other diesel hammers may have continuously adjustable 
207

fuel pumps; for stroke control of such diesel hammers, a reduced combustion 

pressure may be chosen as a percentage of the data file value which corresponds to 

the hammer’s rated energy. However, for construction control, the hammer stroke 

has to be measured, e.g., calculating it from the hammer’s speed of operation in 

blows per minute using a so-called Saximeter, and adjustments of the fuel amount 

have to be made by the operator until the desired, analyzed stroke is achieved.

12.6.4 Vibratory Hammer Considerations

Vibratory hammers come in a variety of models, among them high frequency 

hammers and resonance free hammers. For the high frequency hammers, as they 

are approaching the natural frequency of the piles, it will be difficult to make accurate 

predictions of drivability and or stresses because of the uncertainty of the hammer, 

pile and soil response. Resonance free hammers have a variable eccentric moment 

which would require different hammer data file values for eccentric moments that 

differ from the rated one, should the hammer be run on a reduced setting. However, 

in general, the reduced eccentric moments are only used during the starting and 

stopping cycle so as to reduce the danger of low frequency soil resonance while the 

maximum one is used for the production driving situation.

Vibratory hammers usually perform at 100% efficiency, with the greatest uncertainty 

the actual frequency. Note that often reduced driving times can be achieved with 

lower frequencies than maximum depending on the mass and stiffness of the 

combined soil-pile system.

As pointed out in Example 9, there is a great difference in how the soil responds to a 

vibratory hammer and an impact hammer. In granular, submerged soils the shaft 

resistance may almost complete vanish due to a liquefaction type effect. In cohesive 

soils, if they are sensitive to vibration a similar very pronounced loss of resistance 

may occur during driving. However, in soils where the bond between pile surface 

and soil is not broken, possibly because the hammer amplitude is too small, the 

resistance of the soil attached to pile may lead to refusal. In any event, where the 

vibratory excitation of the soil-pile interface leads to great losses of resistance, the 

relationship between speed of pile penetration rate and soil resistance is virtually 

meaningless while for those soil where the soil-pile bond is not broken, the full soil 

resistance is also not mobilized.
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12.6.5 Batter Pile Considerations

Pile top bending due to poor hammer alignment is more likely to occur in battered 

than in vertical driving because of difficulties with maintaining a good hammer–pile 

alignment. The problem is aggravated during restrike, because realigning hammer 

and pile is then even more difficult than maintaining alignment during initial driving.

Inclined pile driving also causes, sometimes unexpected, bending stress problems if 

the piles are not properly supported and guided by the hammer leads. In fact, for 

offshore leads where the pile supports the non-axial weight component of pile, leads 

and hammer, these bending stresses are predictable. Offshore versions of wave 

equation programs have been developed for those situations that include a routine 

for calculating these bending stresses which are then superimposed to the dynamic 

stresses. However, for piles supported by fixed, swinging or semi-fixed leads it is 

assumed that the piles are supported in such a way that no significant bending 

stresses result along the length of the pile that is supported by the leads. This must 

be confirmed on site by a thorough inspection.

It has been mentioned that external combustion hammers (primarily air or hydraulic 

hammers) possibly do not reach the full effective stroke during inclined or batter pile 

driving, because of limitations of the effective (vertical) ram travel. This should be 

modeled by a reduced stroke. It must be expected that the full rated hammer energy 

will not be available even if the hammer efficiency is satisfactory. The increased 

friction in the hammer due to the pile batter can be estimated under consideration of 

a friction coefficient and batter angle. The wave equation software’s Help section 

should provide some helpful values. For instrumented hammers, the friction loss is 

generally internally measured and/or compensated for. As the pile is driven, energy 

measurements whether in the hammer or by pile top measurements, will be able to 

assess the actual hammer energy transferred from the hammer to the pile. 

For open end diesel hammers, also called internal combustion hammers, the ram 

travel is unlimited and it is likely that the actual, inclined ram travel is greater than 

the vertical ram travel under similar pile and soil conditions. However, friction effects 

should be accounted for by a reduced efficiency value as discussed for the external 

combustion hammers.
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12.6.6 Hammer and Pile Cushion Considerations

Hammer cushions are primarily used to protect the hammer while pile cushions 

protect concrete pile tops. Direct Drive hammers use no cushioning material and 

are sometimes used for steel pile driving situations.

Cushion materials are subjected to destructive stresses during their service and, 

therefore, continuously change properties. For hammer cushions, a variety of manmade materials (e.g., Micarta, Conbest, and Nylon among others) are acceptable 

while wood chips as a hammer cushion are totally unpredictable and therefore 

should never be allowed. Frequently, hammer cushions are engineered as 

sandwiches of aluminum plates (to extract heat) and softer cushioning materials. 

The input values for the combined stiffness of two cushion materials can be easily 

calculated from the individual material properties by remembering that the inverse of 

the combined stiffness, kc, of two springs in series is the sum of the inverse of the 

two individual springs (1/kc = 1/k1 + 1/k2). However, for most commonly used 

cushions the Help section of the wave equation software contains extensive material 

property tables. Wood chips or softwood with the grain parallel to the pile axis as a 

hammer cushion have totally unpredictable material properties and therefore should 

not be allowed.

It has been explained in Example 4 that pile cushions experience a particularly 

pronounced increase in their stiffness during driving, because they are generally 

made of soft wood with its grain perpendicular to the load. Typically, the 

effectiveness of wood cushions in transferring energy increases until they begin to 

burn and quickly deteriorate. This typically happens after approximately 1500 

hammer blows. For conservative stress predictions, the harder, used cushion could 

be modeled by an increased elastic modulus and reduced thickness. In the United 

States two types of pile cushions are most frequently encountered. The most 

common pile cushion consists of plywood sheets and the second most common pile 

cushion is made of oak boards. Improved agreement with measurements can be 

achieved if new and used plywood cushions are analyzed with elastic moduli of 30 

and 75 ksi, respectively (Rausche et al. 2004). For oak board pile cushions, the 

respective recommended values for new and used cushions are 60 and 90 ksi. 

These elastic modulus values should always be used with the nominal 

(uncompressed) thickness. For conservative nominal resistance predictions, a less 

effective pile cushion may be modeled using by a somewhat lower, maybe 50% 

lower than normally recommended, input for both elastic modulus and coefficient of 

restitution.
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Uncushioned or direct-drive diesel and hydraulic hammers are also frequently 

encountered when steel piles are driven. The advantage of these hammers is 

obvious: their energy transfer to the pile is not subject to varying cushion properties. 

For the wave equation analysis, the stiffness of the spring between hammer and 

helmet is derived from the elastic properties of either ram or impact block (diesels)

since there is no hammer cushion. (This stiffness is very high, much higher than the 

stiffness values of most other components within the system, and for numerical 

reasons, may lead to inaccurate stress predictions. Analyses with different numbers 

of pile segments (the automatically selected number of pile segments is the pile 

length divided by 3.3 feet) would show the sensitivity of the numerical solution. In 

general, the greater the number of pile segments, the more accurate the stress 

calculation). Choosing 1 foot long segments (the number of pile segments then 

equals the pile length in feet) would generate a much more detailed pile model than 

is standard.

12.6.7 Selection of Soil and Rock Parameters

The greatest errors in nominal resistance predictions are usually observed when the 

soil resistance has been improperly considered. A very common error is the 

confusion of factored design loads with the wave equation's nominal resistance 

(usually called ultimate capacity in the wave equation documentations). Note that 

the wave equation nominal resistance always must be multiplied by a resistance 

factor which depends on the nominal resistance verification method and this product 

has to be greater than the factored load. In the past, and that is still the way the 

wave equation documentation describes the approach, the ultimate capacity was 

divided by a factor of safety to yield the allowable design load. Resistance factors 

suggested by FHWA and AASHTO are discussed in Chapter 7.

Since the soil is disturbed at the end of driving, it then often has a lower nominal 

resistance (occasionally also a higher one) than at a later time. For this reason, a 

restrike test should be conducted to assess the nominal resistance after time 

dependent soil strength changes have occurred. However, restrike testing is not 

always easy. The hammer is often not warmed up and only slowly starts to deliver 

the expected energy while at the same time the nominal resistance of the soil 

deteriorates. Depending on the sensitivity of the soil, the penetration resistance may 

be taken from the first 3 inches of pile penetration even though this may be 

conservative for some sensitive soils. For high penetration resistances (e.g., more 

than 10 blows/inch) smaller than 3-inch penetrations should be considered or the 

penetration for ten blows should be accurately measured. For sensitive soils with 

quickly lost soil setup resistance, measuring the penetrations of individual blows in 
211

the beginning of a restrike should be attempted if possible. Of course, this 

information then has to be used together with the energy transferred to the pile head 

for each blow during the early restrike.

For construction control, rather than restrike testing many piles, it is more 

reasonable to develop a site specific setup factor in a pre-construction test program. 

As long as the hammer is powerful enough to move the pile during restrike and 

mobilize the soil resistance, restrike tests with dynamic measurements are an 

excellent tool to calculate setup factors. For the production pile installation criterion, 

the required end of driving resistance is then the required nominal resistance divided 

by the setup factor. From the wave equation calculated bearing graph and with the 

reduced end of driving nominal resistance, the required end of driving penetration 

resistance is found.

Although the proper consideration of static resistance at the time of driving or 

restriking is of major importance for accurate results, dynamic soil resistance 

parameters sometimes play an equally important role. In general, shaft damping 

factors on the order of 0.05 s/ft for non-cohesive soils, 0.10 s/ft for silty sands, clayey 

sands, and sandy silts, 0.15 s/ft for cohesive silts and sandy clays, and 0.20 s/ft for 

cohesive soils are typical. For most soils, the toe damping factor is about 0.15 s/ft. 

The above values are general recommendations for initial driving conditions. 

Damping factors have been observed to vary with waiting times after driving. Thus, 

damping factors higher than recommended above and in the GRLWEAP Manual 

(say twice as high) may have to be chosen for analyses modeling restrike situations. 

Studies on this subject are still continuing. In any event, damping factors are not a 

constant for a given soil type. For soft soils, damping factors may be much higher 

than recommended, and on hard rock they may be much lower. However, choosing 

too low a damping factor may produce non-conservative nominal resistance 

predictions. The program recommendations for damping factors are averages which 

work reasonable well as a first assumption (see input description in Section 12.7). 

After piles have been driven and dynamically tested, appropriate site specific values 

should be available.

Shaft quakes are usually satisfactory as recommended at 0.10 inch. However, for 

displacement piles, toe quakes can vary widely and reach values well in excess of 

the program recommended range of 1/60 and 1/120 of pile diameter or pile width, 

particularly when the soil is saturated and/or rather sensitive to dynamic effects. 

Only dynamic measurements can reveal more accurate soil quakes. However, short 

of such measurements, conservative assumptions must sometimes be made to 

protect against unforeseen problems. Fortunately, toe quakes have a relatively 
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insignificant effect on the wave equation results of piles having most of their 

resistance acting along the shaft. For piles achieving their nominal resistance 

primarily through toe resistance, large toe quakes often develop during driving in 

saturated soils causing the toe resistance to build up only very slowly during the 

hammer blow. As a consequence, at the first instant of stress wave arrival at the 

pile toe, little resistance exists and damaging tension stress reflections can develop 

in concrete piles even if the penetration resistance is high. At the same time, large 

toe quakes dissipate an unusually large amount of energy and therefore cause high 

penetration resistances. Thus, more cushioning or lower hammer strokes may not 

be a possible alternative for stress reductions. Instead, in extreme cases, hammers 

with heavier rams and lower strokes should be chosen to reduce the detrimental 

effects of large toe quakes. Example 6 in Section 12.5.6 illustrates the effect of a 

large toe quake.

Stress predictions, particularly tension stresses, are also sensitive to the input of the 

resistance distribution and to the percentage of toe resistance. If the soil resistance 

distribution is based on a static analysis, chances are that the shaft resistance is set 

too high because of the loss of shaft resistance during driving. It is therefore 

recommended that drivability analyses be performed with shaft resistances reduced 

by estimated setup factors, which will adjust the statically calculated nominal 

resistance to match the conditions occurring during driving.

Soft rock is generally modeled like a soil with quakes and damping values chosen as 

for the underlying material (e.g. clay for a claystone or shale). However, pile toe 

compression stresses can be damaging when the piles are driven to a hard rock. 

Since piles generally do not penetrate into a hard rock and since the plastic stage 

cannot be reached, the main concern is not with the nominal resistance but with the 

pile toe stresses. The resistance mobilized is usually a lower bound conservative 

value. Hence, to be conservative it is recommended to analyze these hard rock 

cases with a toe quake of 0.04 inches. A lower than the normal 0.15 s/ft damping 

value may also be used (e.g., 0.05 s/ft).

Residual stress wave equation analyses are superior to normal analyses in basic 

concept and probably also in results. Unfortunately, because of lack of correlation 

work, no empirically determined dynamic soil constants (quake and damping values) 

can be recommended for use with residual stress analyses. Experience exists for 

fluted and tapered piles which supports using the residual stress analysis with 

commonly used damping factors. Residual stress analyses should be performed 

(maybe in addition to standard analyses for long slender piles with significant shaft 

resistance components,) to assess potentially damaging stress conditions and the 
213

possibility of nominal resistance values which could be much higher than indicated 

by the standard wave equation analysis. Note that residual stress analyses may not 

be meaningful for representation of early restrike situations in which energies 

increase from blow to blow while, in sensitive soils, capacities successively 

decrease. The residual stress analysis assumes that hammer energy and pile 

nominal resistance are constant under several hammer blows. For further 

information please see also Section 12.7.1 and the background information of the 

wave equation computer program.

12.6.8 Pile Modeling Considerations 

In general, the pile is represented more accurately and reliably than other parts of 

the wave equation model. Of course, unit weight and elastic modulus have to be 

well known which is sometimes challenging for concrete and timber. Note that for 

concrete and timber there is a significant difference between the static elastic 

modulus and the dynamic one, the latter usually being significantly higher than the 

static one. Since approximately 2012, it has also been observed that even the steel 

elastic modulus can be higher by up to 7% for large diameter pipe piles than 

normally assumed. Plastic piles are also some times analyzed and because of the 

large variety of materials and their basically non-elastic behavior, dynamic testing is 

needed to determine the material properties prior to performing the analysis. 

Modern wave equation programs automatically generate a pile model which is 

generally satisfactory for most commonly encountered situations. However, there 

are differences in the basic approach between different software products and the 

user should be aware of any limitations of these products in certain situations.

For the Smith-type programs which use mass and stiffness to represent the elastic 

properties of the hammer, driving system and pile, non-linear system components 

like cushions or splices with slacks are relatively easily and quite realistically 

represented. In addition, the default 3.3 foot long pile sections are generally 

adequate although smaller segments can lead to an improved program 

performance. Situations where smaller pile segments or also smaller computational 

time increments are helpful may be highly non-uniform piles or uncushioned driving 

systems. Like the length of the pile segments, computational time increments are 

normally set by the computer program, but can be modified with a factor which 

E.A.L. Smith called a safety factor against numerical instability. This factor is 

normally 1.6 in GRLWEAP and may be increased to 3 or even 5 if unusual situations 

cause the program to give a numerical instability warning.
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For programs which use the characteristics approach to solving the underlying 

differential equation (e.g. Middendorp 2004) the computational time increment must 

be equal to the wave travel time of every segment. This leads to segments of 

different lengths when the pile consists of different materials such as a concrete pile 

with steel stinger. While the characteristics solution has the advantage of truthfully 

modeling ideal elastic situations, such as a square input pulse without degradation, 

there is a natural tendency for waves propagating along materials to maintain lower 

frequency components and shed higher frequency ones. This leads to a smoothing 

of the pulse traveling along the pile which is similar to the degradation observed 

when analyzing with a lumped mass model. Modeling piles with slacks in splices is 

also more difficult with the characteristics method than using the more forgiving 

Smith approach. 

Unusual situations which require special attention are: installations which include 

followers, particularly when a soft cushion separates a steel follower from a concrete 

pile, mandrel driven piles, and composite piles such as a concrete pile with a steel 

stinger. These conditions require modeling using splices with compression and 

tension slacks, parallel piles and more than one toe model. Examples of these more 

complex situations are available in the help section of the GRLWEAP software. 

Another problem frequently encountered is the uncertainty on how to model open 

end pipe piles. Very large diameters (say greater than 120 inches diameter) will 

never plug during driving and may experience some internal friction. This can be 

modeled by using an increased perimeter as discussed in Section 12.5.8. Smaller 

diameter (say 30 inches or less) open ended pipe piles, will generally plug once 

driven a sufficient, diameter dependent distance into a dense or very dense material. 

Intermediate pile sizes may or may not plug and then behave either like the piles 

with larger or the smaller diameters. The pile model may have to include a 

consideration of the mass effect and soil model needs to address the effective toe 

area over which toe resistance acts. While this is primarily a soil model problem, it 

potentially affects the mass properties of the pile model over the plug length. As 

noted in Chapter 7, Holloway and Beddard (1995) also reported that hammer blow 

intensity (impact force and energy) influenced plug formation and slippage. Brown 

and Thompson (2015) published a Synthesis of the current practice of design and 

analysis of large open ended pipe piles which provides further insight on this issue. 

When the pile model to be analyzed is unusual in its complexity, it is recommended 

to perform sensitivity studies by comparing results with different numbers of pile 

segments and/or reduced time increments and possible maximum and minimum 

values of pile properties.
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12.6.9 Comparison with Dynamic Measurements

Often, wave equation predicted stresses and nominal resistance values initially 

appear to agree quite well with results from field dynamic measurements, described 

in Chapter 10. However, there are additional observations and measurements that 

should be compared, such as stroke or bounce chamber pressure and transferred 

energy. Often transferred energy values are somewhat lower than calculated, and 

adjustment of hammer efficiency alone may improve energy agreement but produce 

problems with driving stress and nominal resistance agreement. Thus, instead of 

adjusting hammer efficiency, the cushion stiffness or coefficient of restitution may 

require reduction. Sometimes matching of measured values can be very frustrating 

and difficult, and the task should be done with reason. Matching stresses and 

transferred energies within 10% of the observed or measured quantities may be 

accurate enough.

Note: The wave equation maximum stresses in the final summary table can occur 

anywhere along the length of the pile and therefore at a location different from where 

the field measurements were taken. It is therefore important to check the maximum 

driving stresses in the Extrema Tables for the pile segment that corresponds to the 

measurement location when comparing wave equation calculation and field 

measurement results.

Matching wave equation results to field observations and measurements is often 

referred to as a Refined Wave Equation Analysis (Rausche et al. 2009). The 

following procedure requires that wave equation input parameters for hammer, 

driving system, and soil resistance are adjusted and then wave equation analyses 

are performed for the field verified nominal resistance. The following data 

preparation steps and successive input parameter adjustments generally lead to an 

acceptable solution. As noted in Section 12.6.8, multiple wave equation software 

programs exist and some use different models. The correlation procedure below is 

specifically based on the hammer, pile, and soil models used in GRLWEAP and 

CAPWAP. Procedures for other wave equation and/or signal matching programs 

may differ depending upon the hammer, pile, and soil models used in those 

programs.

a. Set up a table with the observed stroke or bounce chamber pressure for 

diesel hammers, and measured values of compression stresses and 

transferred energy, both at the measurement location. Include in this table for 

concrete piles the PDA calculated maximum tension stresses. These values 

should be averages over several consistent blows of pile installation or the 
216

earliest consistent blows of restrike testing. Additional matching quantities 

are the CAPWAP calculated nominal resistance and penetration resistance.

b. Set up the GRLWEAP wave equation model to run bearing graphs for the 

actual hammer, pile, and driving system with total nominal resistance, 

resistance distribution, quake, and damping obtained from dynamic field 

measurements and by signal matching.

c. Perform wave equation analyses and compare results with table values from 

step a. Adjust hammer efficiency (for diesel hammers, also maximum 

combustion pressure) until agreement between measured and wave equation 

computed compression stress and transferred energy (for diesel hammers, 

also stroke) is within 10%. For steel piles modifications of the hammer 

cushion stiffness and/or its coefficient of restitution and for concrete piles 

adjustments of the pile cushion stiffness and/or its coefficient of restitution 

may also be needed.

d. After an initial agreement has been achieved for transferred energy and pile 

top compression stress, compare calculated penetration resistance for 

CAPWAP nominal resistance and associated maximum tension stresses. For 

steel piles, adjust hammer cushion stiffness and coefficient of restitution, and 

for concrete piles, adjust the equivalent pile cushion parameters, together 

with efficiency, to improve agreement of penetration resistance and tension 

stresses within the 10% tolerance.

e. Adjust the hammer efficiency to values not greater than 0.95 and not less 

than 50% of the standard recommended hammer efficiency values for that 

hammer type. The exceptions are hammers whose stroke input is based on 

measured impact velocity, therefore efficiency values greater than 0.95 are 

possible. Adjust cushion coefficients of restitution between 0.10 and 1.0.

f. If penetration resistance and stresses cannot be simultaneously matched by 

adjusting hammer and driving system parameters, change the shaft and toe 

damping and the toe quake simultaneously and proportionately to achieve 

agreement between measured and computed penetration resistance. Under 

certain conditions, it may also be necessary to change the wave equation 

damping model from Smith to Smith-Viscous. 
217

Perfect agreement should not be expected between wave equation results and 

quantities derived from field measurements. The reason is primarily a difference 

between the measured pile top force and velocity and the corresponding quantities 

obtained by the wave equation driving system model. Also, there may be some 

differences in the pile model and soil models between the different analyses 

programs used for wave equation and signal matching. Plots of wave equation 

calculated and dynamic test measurements such as force and velocity can be easily 

generated and can sometimes explain the differences between observed or 

measured and calculated values.

12.7 WAVE EQUATION ANALYSIS INPUT

As described in the previous sections, the input for a wave equation analysis 

consists of information about the soil, pile, hammer, cushions, helmet, splices, and 

any other devices which participate in the transfer of energy from hammer to soil. 

This input information is usually gathered from contract plans, the contractor's 

completed Pile and Driving Equipment Data Form (Figure 12-24), soil boring, and a 

static pile nominal resistance analysis. In a case where the contractor proposes 

using a follower as part of the driving system, detailed drawings of the follower 

should also be obtained. Helpful information can also be found in the "Help" display 

of the wave equation program. These tables are correct only for ideal situations but 

may yield valuable data before a specific driving system has been identified. In 

general, contractors tend to assemble equipment from a variety of sources, not all of 

them of a standard type. It is therefore important to check and confirm the 

equipment that the contractor has actually included in the driving system on the 

project.
218

Figure 12-24 Pile driving and equipment data form. 
219

The following sections explain the most important input quantities needed to run the 

GRLWEAP program. For a more detailed explanation of input quantities, reference 

is made to the program's Help Section (function key F1 or F3 or click on Help).

The (second topic) of the Help Menu (F1 or click on Help) explains the Main Input 

screen and all of its menus, data entry fields, and information indicators. Figure 

12-25 shows this Help Window as it first appears and subdivides the Main Input

screen into 10 major sections:

A. Standard Window Menus

B. Icons for standard Windows Operations

C. Icons for GRLWEAP displays and operations

D. GRLWEAP Drop-Down Menus

E. Input fields for Title and Hammer Selection

F. Hammer Parameters and Pile Material Selection

G. Hammer and Pile Cushion Input

H. Pile Data Input

I. Ultimate Capacity (Nominal Resistance) Input or Gain/Loss Factors

J. Soil Parameters

Although a simple bearing graph analysis only requires input in the GRLWEAP Main 

Input Screen, it is recommended to utilize the step-by-step input requests generated 

after clicking on the “New Document” icon (or New in the File Menu).

The Job Information window shown in Figure 12-26 will display first, accepting input 

of a title of up to 40 characters and the assignment of a file name and directory. 

Browse may be used to navigate the user’s computer and assign the desired 

directory.

Clicking on Next will open up the Select Hammer window shown in Figure 12-27. 

The GRLWEAP program includes a hammer data file in which the major mechanical 

properties of approximately 1000 hammers are stored. By selecting an identification 

number (ID) and/or corresponding hammer name in the List of Hammers window, 

the user prompts the program to automatically input the selected hammer’s 

properties. Note that the automatic hammer input assumes use of a well maintained 

and unmodified hammer.
220

Figure 12-25 GRLWEAP help window for main input form. 

While initially all hammers are displayed in the order of hammer ID number, the 

display may be reorganized by certain hammer types or manufacturers. Hammer 

types are OED (Open End Diesels), CED (Closed End Diesels), ECH (External 

Combustion Hammers, including the air, steam, hydraulic and drop hammer 

categories), and VIB (Vibratory Hammers). The user can also organize the contents 

in the List of Hammers window by hammer Name, Type, Ram Weight or Rated 

Energy by clicking on the column heading.
221

Figure 12-26 Job information window. 

Figure 12-27 Select hammer window. 
222

The next section involves the Analysis Type window, displayed in Figure 12-28. 

For a simple Bearing Graph, the Proportional Shaft Resistance option is the 

default. It assumes constant percentages of shaft resistance and end bearing for all 

nominal resistance values to be analyzed. The alternate bearing graph options 

analyze the various nominal resistance values either assuming a Constant Shaft 

Resistance or a Constant End Bearing. 

A modified Bearing Graph approach, the Inspector’s Chart provides the possibility 

of analysis with an increasing stroke (or hammer energy values) for a single nominal 

resistance value. This option is useful for diesel hammers, whose stroke can vary 

and/or be adjusted by different fuel settings, and for hydraulic hammers, whose 

energy level can be selected on the hammers’ control panel.

The user may also choose the Drivability option. It requires as an input the unit 

shaft resistance and unit toe resistance as a function of pile penetration and, 

therefore, requires an accurate static soil analysis. The resulting output will show 

the corresponding nominal resistance values together with calculated penetration 

resistance (blow count), pile stress maxima, and other quantities and, thus, indicates 

the complete, expected driving behavior.

Figure 12-28 Analysis type window. 
223

Clicking on Next brings up the Pile Input window, illustrated in Figure 12-29. After 

selection of the pile material, i.e., Concrete, Steel, or Timber, the program inputs 

default values for pile top elastic modulus, coefficient of restitution, and specific 

weight in the corresponding fields and also, for concrete pile material, activates the 

pile cushion input section. As with the hammer cushion, described below, the user 

may utilize the Area Calculator and the Cushion Material Properties Help by 

pressing the F3 function key or directly input a stiffness. Additionally, selection of 

the pile material will automatically select the pile damping parameter which is 

accessible through Options, General Options, Damping. The user may adjust the 

aforementioned defaults but must enter the initial inputs for the Pile Length and 

cross Section Area. For the latter, the user may again employ the Area Calculator, 

shown in Figure 12-30, which also provides the Pile Size, Perimeter and Toe Area 

based on pile type and pile dimensional information. These quantities are 

particularly important for the drivability analysis when calculating the nominal 

resistance from the unit resistance values.

Figure 12-29 Pile input window. 

It is important to note that for non-uniform piles the input quantities of Cross 

Sectional Area, Elastic Modulus, Specific Weight and Perimeter, in this window 

only refer to the pile top. Also Toe Area and Pile Size may need correction for nonuniform piles, the latter because it is used for calculating a recommended toe quake.
224

Once the data entry wizard has been finished, the non-uniform quantities must be 

entered in the P1 window, accessible after clicking on the pile type drop-down menu. 

As form most default values, those automatically chosen for pile elastic modulus and 

specific weight may or may not be correct and must be reviewed by the program 

user. For example, for concrete or timber piles, measurements could indicate other 

values. Pressing F3 with the cursor on the Elastic Modulus or Specific Weight input 

field brings up added Help information. The following information is required 

information.

Length is the total pile length in the leads in feet. For example, if plans 

require a pile of 50 feet in length but the contractor is driving 60 long piles, the 

proper analysis length would be the full 60 feet. If pile sections are spliced 

together to form a longer pile, an analysis before and after splicing may be of 

interest. In such cases, the Length may be either the length of a single 

section before splicing or the combined length after splicing.

Penetration refers to the analyzed pile toe penetration below grade in feet for 

Bearing Graph or Inspector’s Chart and final penetration for Drivability 

Analyses. This measurement must use the same soil grade reference as that 

of the soil resistance distribution.

Section Area is the pile cross section area at the pile head in inch2. 

Elastic Modulus is the elastic modulus of the pile material at the pile head in 

ksi. 

Spec Weight is the weight per unit volume of the pile material at the pile 

head in lbs/ft3

. This value is used for calculating the mass of the pile material 

by division with 32.17 ft/s2. The program provides for a modification of the 

pile’s weight component by modification of the gravity acceleration value in 

Options/General Options/Numeric to reflect, for example, the effect of 

buoyancy or pile inclination on the pile weight on the soil.
225

Figure 12-30 Area calculator window

After the pile input is done, clicking on Next will open up the Hammer Cushion 

window, as shown in Figure 12-31. GRLWEAP offers an extensive data file for 

those situations in which the contractor’s available equipment is unknown. The data 

file has been made possible courtesy of the various manufacturers and dealers 

whose products are listed. Please note that this file is neither complete nor 

necessarily appropriate for all situations, as the contractor may not follow the 

manufacturer’s recommendations. The required information consists of:

Area is the area of the hammer cushion perpendicular to the load in inch2. 

Elastic Modulus is the elastic modulus of the hammer cushion material in 

ksi. 

Thickness of the hammer cushion. For sandwiched cushions, this is the 

thickness of the cushion material that corresponds to the elastic modulus in 

inches. If the entire stack thickness is entered, the combined elastic modulus 

of the sandwich and the striker plate is not included. If no hammer cushion 

exists, leave this value and the stiffness value at zero.

C.O.R. is the Coefficient of Restitution of the hammer cushion material.

Stiffness of the hammer cushion in kips/inch. Use of this optional input will 

override the inputs for area, elastic modulus, and thickness.
226

Helmet Weight, consisting of the combined weight of the helmet, hammer 

cushion, striker plate, inserts, and all other components located between the 

hammer and pile in kips. The input may be zero if there is no helmet mass.

Figure 12-31 Hammer cushion window. 

Ideally, the contractor would provide the above drive system data for his actual 

hammer system. However, if not available, the required hammer cushion data may 

be selected using one of three different methods:

1. The hammer cushion Stiffness and Coefficient of Restitution may be known 

from other analyses and can be input directly into the appropriate fields. In 

such cases, hammer cushion area, elastic modulus, and thickness are not 

needed.

2. If some or all of the driving system data is to be retrieved from the program 

data file, merely pressing F3 while the cursor is on one of the associated input 

fields and then clicking on Manufacturer’s Recommended Driving System 

opens a listing of the recommended input. The user may transfer the

suggestions in whole or part to the input sheet.

3. If the cushion material area, thickness, and type are known but modulus of 

elasticity and Coefficient of Restitution are not, pressing F3 while the cursor is 
227

on the elastic modulus field and then selecting Cushion Material Properties 

brings up a list of frequently used cushion materials and their properties. 

These values can be transferred directly to the hammer cushion data input 

fields.

The Next input sections for bearing graph or inspector’s chart analyses may be done 

in the dynamic soil parameters window on the Main Input Form or Soil Profile 

Input window, displayed in Figure 12-32. (For Drivability analyses, the S1 Form is 

opened as later discussed.) The most convenient input is through the ST analysis in 

the Soil Profile Input window. There, the user first specifies the:

Number of Soil Layers. It is recommended to divide the soil into layers of not more 

than 10 feet in thickness for improved accuracy.

Final Penetration Depth is the distance from grade to that depth to which data is to 

be given in feet. The window will at first display the value entered under the pile 

information. However, it may be changed here with the exception that it cannot be 

greater than the pile length.

Water Table is the distance from grade in feet where the water table begins. If 

grade is underwater, enter zero or a negative value; the later will show on the Main 

Screen graphics the water depth. As far as buoyancy is concerned both zero and 

negative values give the same result.

Effective Overburden at Grade is the intensity of any overburden pressure in ksf. 

For example, in the case of an excavation of limited extent (trench), the depth of 

excavation times the soil unit weight equals the effective overburden.

For each layer, the analyst then enters:

Either the Layer Bottom Depth or the Layer Thickness in feet.

The layer soil type as either Granular (non-cohesive soil for primarily sandy or other 

coarse grained soils) or Cohesive (for clays and silts) and selects as sub types the 

density or consistency of the layer. For intermediate soil types or non-cohesive silts, 

it may be conservative to choose “cohesive”, since soil damping is then assigned a 

higher value. However, under all circumstances, the analyst should review the 

results obtained from this very simplified analysis.
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Figure 12-32 (a) Soil profile input window for soil type based static soil analysis and 

(b) Soil parameter input window for bearing graph analysis. 

After clicking Update, the program will display a nominal resistance (Ru) and a 

nominal shaft resistance (Rs) value. These two results pertain to the Final 

penetration Depth, where the ratio Rs/Ru is the percentage of shaft resistance and 

one of the soil resistance inputs generated by the routine. Under no circumstances 

should these values be used for pile design purposes. The results are based on the 

following two methods:

For Non-Cohesive Soils

Using the Effective Stress Method, the unit shaft resistance is fs= β σ’v, with β being 

the Bjerrum-Burland beta coefficient as tabulated in Table 12-8 and σ’v being the 

effective vertical stress at the midpoint of a given soil layer. The unit toe resistance 

is qp = Nt σ’vt, where Nt is a toe bearing capacity coefficient (see Table 12-8) and σvt

is the effective overburden pressure at the pile toe. Both fs and qp are subjected to 

certain specified limits.

For Cohesive Soils

For cohesive soils, ST applies a modified α-method, also called the total stress 

method, and relies on the unconfined compression strength (qu) of the soil layer. 

The qu-value and, based on it, the unit shaft and unit toe resistance values are 

shown as a function of both soil type and a representative N-value in Table 12-9. 
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Table 12-8 Soil Parameters in ST Analysis for Granular Soil Types

Soil Type SPT N

Friction 

Angle 

φ

Unit 

Weight 

pcf

β Nt

Max. Unit 

Shaft Res. 

ksf 

Max. Unit 

End Bearing 

ksf 

Very Loose 2 25-30 86 0.203 12.1 0.5 50

Loose 7 27-32 102 0.242 18.1 1.0 100

Medium 20 30-35 118 0.313 33.2 1.5 150

Dense 40 35-40 125 0.483 86.0 2.0 200

Very Dense 50+ 38-43 141 0.627 147.0 4.o 400

Table 12-9 Soil Parameters in ST Analysis for Cohesive Soil Types

Soil 

Type SPT N

Unconfined 

Compression

Strength 

ksf 

Unit 

Weight 

pcf

Max. Unit 

Shaft Res. 

ksf 

Max. Unit 

End Bearing 

ksf 

Very Soft 1 0.25 111 0.07 1.1

Soft 3 0.75 111 0.23 3.3

Medium 6 1.50 118 0.40 6.7

Stiff 12 3.00 131 0.81 14.0 

Very Stiff 24 6.00 131 1.30 27.0 

Hard 32+ 8.00+ 121-140 1.60 36.0 

After the soil types of all layers have been entered, the program computes the 

percentage and distribution of shaft resistance, the average shaft damping 

parameter, and the toe quake. These wave equation input values are based on pile 

penetration, water table depth, pile size, pile perimeter, and pile toe area. Damping 

and quake values are adjusted considering soil and pile type. Again, the analyst is 

responsible for checking these values and should be aware that this analysis is not 

applicable to non-uniform piles.

It is very important that the user carefully reviews the wave equation input 

parameters resulting from this very simplified static soil analysis, possible in the Soil 

Parameters Input window (see Figure 12-32b). Particular attention should be paid 

to the pile toe area because the shaft resistance percentage and toe quake directly 

depend on its magnitude. Also, it is recommended to perform comparative 

analyses, for example, when the soil type does not clearly fall into either the 
230

cohesive or granular categories. In such cases, results for both soil types should be 

obtained and compared. The ST generated input parameters should be reviewed 

once the input wizard has been finished and the main screen is displayed.

Help pertaining to both soil type input and soil quakes and damping appears in the 

program Help Menu under GRLWEAP Input Forms, Soil Type-Based Input Form 

and GRLWEAP Component Parameters, Soil Parameters, respectively. It is also 

recommended that the user carefully review both the PDI (2010) Background Report 

and the program Help.

For Bearing Graphs or Inspector’s Charts, the user must input between one and ten 

nominal resistance values in the Ultimate Capacity window shown in Figure 12-33. 

Several options are available including values spaced at constant increments (Incr.), 

generated by pressing “Interpolate” to interpolate between the first and last entries, 

and Automatic Capacities, based on the pile cross section properties. It is 

recommended to analyze nominal resistance values that will provide a meaningful 

bearing graph for both easy and hard driving conditions. The input wizard is now 

finished. The completed Main Input screen should resemble that shown in Figure 

12-39. To perform a more complex analysis, additional inputs may be made by 

specifying a Non Uniform Pile or a more detailed soil resistance distribution in 

Variable Resistance Distribution.

Figure 12-33 Window for entering up to 10 nominal resistance values for bearing 

graph and inspector’s chart analyses.
231

For Drivability, instead of nominal resistance values, the analyst must input 

Resistance Gain/Loss Factors. Figure 12-34 shows the related window. The 

analyst may perform at most 5 analyses at each specified depth and provide at most 

five associated gain/loss factors for both the pile shaft and toe. These factors are 

related to the soil resistance parameters to be entered in the S1 Form (Figure 12-35) 

discussed below. A factor of 1.0 implies no change in soil strength during driving 

and thus that no resistance gain or loss will be analyzed. A factor less than 1.0 

proportionally reduces the resistance values under consideration of their relative 

setup factors and thus reflects that the soil resistance is lower during driving and 

increases after pile installation, i.e., soil setup. A factor greater than 1.0 

proportionally increases the resistance values and thus reflects the soil relaxation 

scenario, i.e., where the soil resistance is greatest during driving. In most cases, it is 

sufficient to enter two values for the shaft analysis. The first shaft value, marked 1, 

would be the inverse of the highest soil setup factor entered in the S1 Form and 

would represent the greatest resistance loss during driving along the shaft. The 

associated toe resistance factor, Toe 1, would be set to 1.0 to indicate neither gain 

nor loss of toe resistance during driving. For the second analysis, Shaft 2 and the 

associated factor Toe 2 would be set to 1.0. This latter input then reflects the 

absence of both gain and loss during driving at each depth analyzed (see Figure 

12-34).
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Figure 12-34 Resistance gain/loss factors in main input form for drivability analysis. 

Next for Drivability analyses, the GRLWEAP program requires input in the S1 Form 

(Figure 12-35). Important inputs for each soil layer are (refer also to descriptions for 

the equivalent bearing graph inputs):
233

Depth is the soil layer distance below grade or mudline in feet.

Nominal Unit Shaft Resistance in ksf is determined by a static geotechnical 

analysis (e.g., the GRLWEAP SA routine). GRLWEAP multiplies this input by 

the pile perimeter, the segment length, and a soil layer specific gain/loss 

factor to yield the shaft resistance at the segment.

Nominal Unit Toe Resistance in ksf equals the unit toe resistance 

determined by a static geotechnical analysis.

Skin quake is the shaft quake in inch, usually left at the default value of 0.10 

inch.

Toe quake in inch is per the Soil Parameter Help.

Skin damping is the shaft damping in s/ft as per the Soil Parameter Help.

Toe damping in s/ft, usually left at the default value of 0.15 s/ft.

Setup factor is based on site specific knowledge and, in conjunction with the 

resistance gain/loss factor, determines for each soil layer the soil resistance 

to driving.

The parameters of Limit Distance and Setup Time allow for a qualitative 

evaluation of soil strength change during driving interruptions, providing for 

more detailed analyses of splice time interruptions. These parameters have 

no influence on results as long as entered Wait Times in the Depths, 

Modifiers Input Form (see Figure 12-38) are zero.

Figure 12-35 S1 form for soil resistance vs. depth input.
234

Should the user utilize the static analysis (SA) routine that is built into the 

GRLWEAP program to complete the S1 Form, clicking the SA icon will open the 

Figure 12-36 window. First, selection of Profile and New allows specification of the 

following quantities in the Static Analysis General Information window (refer to 

Figure 12-36):

The Total Number of Soil Layers to be included between grade and a depth 

equal to the total pile length in feet.

The depth of the Water Table in feet.

Overburden Pressure in ksf; see also ST above for an explanation of these 

inputs.

After closing the Profile window, the soil layer specific input can be made in the SA

window. The following information should be provided (see Figure 12-37):

Layer Bottom Depth or Layer Thickness is in ft.

If the SPT N-value is known, choose from Gravel, Sand (with sub types 

indicating Grading and Grain Size), Silt, Clay, or Rock. Then enter the SPT 

N-value (not greater than 60), and the program will calculate a unit resistance 

and a unit weight for the soil layer.

If the SPT N-value is unknown, choose Other and either Cohesive or 

Cohesionless and then provide the Unit Weight in kips/ft3 and the Unit Shaft 

Resistance and Unit Toe Resistance, both in ksf. The program will reduce 

the input unit weight value below the water table to yield an effective 

overburden.

Upon user request, the SA routine will also fill in the Other Parameters, i.e., 

the input values for quake, damping, setup factor, limit distance and setup 

time columns. Oftentimes these values do not differ from the default values

specified earlier and then can be left at zero. Again, the user should carefully 

review the automatically generated input values prior to performing the actual 

wave equation analysis.

The SA routine is basically an effective stress method with different 

approaches for sand, silt and clay. This method is described in detail in the 

Background Report of the GRLWEAP program (PDI, 2010) and that 
235

description should be reviewed prior to using this approach. It applies very 

conservative limits on the unit resistance values which may be nonconservative as far as drivability analyses is concerned (calculated driving 

resistances and stresses may be low). Also the method is only applicable to 

piles with straight sections (not applicable to tapered piles) and should never 

serve as the sole static soil analysis method for a pile design. In fact, it is 

always prudent to compare several static analysis methods for an 

assessment of the range of possible results.

Figure 12-36 New profile window for the SA static analysis. 
236

Figure 12-37 Soil property input window for the SA static analysis. 

In a Drivability analysis, another required input is found in the Depths, Modifiers 

Input Form, accessible after clicking the D icon (see Figure 12-38). This form 

provides for the input of:

Depth to be analyzed in feet is a required input for at least two different depth 

values.

The pile Temporary Length in feet may be less than or equal to the length 

value given as the final length input and allows for consideration of a reduced

pile length prior to splicing.

Wait Time in hours, which would be applicable if driving were to be 

interrupted, for example, for splicing operations. This input is only useful for a 

qualitative assessment of setup effects during the driving interruption, which, 

in turn, is a function of the Limit Distance and the Setup Time. Also it only 

applies to the first gain/loss factor. (As mentioned, this is rarely needed for 

highway construction projects.)

Stroke and Efficiency allow for variation of these hammer parameters as a 

function of depth. If not specified, the values input previously will be 

considered.
237

Diesel Pressure input allows for a modification of the hammer setting and/or 

stroke.

Pile Cushion Coefficient Of Restitution (COR) or Stiffness Factor can 

also be varied as a function of depth. For example, and considering the 

recommendations for new and used pile cushion parameters, the Stiffness 

Factor may be gradually increased from 1.0 to 2.5 if the cushion elastic 

modulus was specified earlier with the “New” elastic modulus.

Figure 12-38 Depths, modifiers input form. 

12.7.1 Other Analysis Options

A variety of options exist in the GRLWEAP program for non-standard input, 

analyses, and output. The setting of important options is indicated on the Main Input 

screen (see Figure 12-39). Please refer to the program Help Menu for additional, 

less frequently used options. 

Important options pertain to the modification of certain hammer parameter, pile 

model, and soil resistance input. These options are generally accessible by clicking 
238

Options/General Options, or Hammer Parameters, or Pile Parameters, or Soil 

Parameters. For proper hammer modeling, the Efficiency and Stroke values 

contained in the hammer data file must often be modified. This modification can be 

done on the Main Input screen below the Hammer Information window. 

Alternatively, these and other hammer details may be modified by clicking on 

Options, Hammer Parameters. Relevant quantities are explained in the Help Menu 

and will not be further discussed here.

Figure 12-39 Completed main input for a simple bearing graph. 

As mentioned, efficiency is a very important hammer parameter. The efficiency 

values in the hammer data file were selected according to the observed average 

behavior of all hammers of the same type. However, depending on a particular 

hammer’s make or state of maintenance, the hammer may perform differently than 

assumed, and its parameters should be adjusted accordingly. Furthermore, 

because of uncertainties in actual hammer performance, greater and lesser 

efficiency values should be analyzed for conservative stresses and blow counts, 

respectively. Finally, efficiency should be adjusted for an inclined or batter pile 

which is simplified in the Options/Pile Parameters/Batter Inclination Input Window 

shown in Figure 12-40. Refer to the Help Menu for recommended efficiency 

reductions and further guidance for inclined pile driving.
239

Figure 12-40 Batter/Inclination input window.

Stroke in feet is a useful performance parameter for single acting diesel hammers 

whose ram is visible and whose stroke is not equal to the rated value. For other 

hammers, energy level may be known, but since stroke is equal to energy divided by 

ram weight, stroke serves as an input for an adjusted hammer energy setting.

Pressure in psi is important for diesel hammers when the calculated and observed 

hammer strokes differ. A new pressure value may be tried for better agreement. 

Also, if the hammer is physically run at a reduced fuel setting, the pressure value 

should be reduced in the program accordingly.

Several options for Pile Parameters are available in GRLWEAP primarily for the 

purpose of flexibility in pile model generation. The status of many of these options is 

indicated under the Pile top Information on the Main Input Form.

The Number of Pile Segments may either be automatically set based on segment 

lengths of 3.3 feet, or the user may use a different number by clicking on Options, 

Pile Parameters, Pile Segment Option. Usually the program default of 3.3 feet 

segment lengths yields satisfactory accuracy. To avoid loss of this computational 

accuracy, only segments smaller than the default value should be entered. In the 

Pile Segment Option, the user can also modify the relative length of the segments 
240

(the information field marked S-Length would then be set to Man. for manual) and 

enter the segment stiffness and mass values. In the latter case, the information field 

marked S-ST, Wt would be set to manual.

Some piles are spliced with devices that allow for slippage during extension or 

compression. In Options/Pile Parameters/Splices, the user can choose the Number 

of Splices to be modeled and, after clicking on Update, edit the entry fields shown in 

Figure 12-41. For each splice, the user enters the Distance in feet of the splice 

location referenced from the pile head, the Tension Slack in feet, i.e. the distance 

that the splice can extend without transmitting a tension force, the C.O.R. or 

coefficient of restitution for the spring representing the spliced section, and the 

Compression Slack in feet, i.e., the distance that the splice can compress with a 

spring stiffness which increases linearly from zero to its loading value. The Main 

Input Form indicates the selected number of splices; the graphic of the hammer, 

driving system, pile, and soil model on the Main Input Form also indicates a splice 

with a slight gap in the pile representation. Note that neither an uncracked welded 

splice of a steel pile nor a well performing epoxy splice of a concrete pile requires 

slack modeling. These splices do not allow for slippage and, therefore, should be 

modeled as a uniform pile section and not as a splice with a slack.

Figure 12-41 Slack/Splice information input window. 
241

Non-uniform pile properties are modeled by selecting Non-Uniform Pile from the Pile 

Type drop-down menu which activates the P1 window. The user should complete 

the necessary information by specifying pile variations and adding the necessary 

number of rows immediately above and below a change of cross sectional area (XArea), elastic modulus (E. Modulus), specific weight (Spec. Wt.), Perimeter, and 

Critical Index (Crtcl. Index) (see Figure 12-42). Note that Perimeter is needed for 

the computation of total shaft resistance in Drivability analyses. The Critical Index 

input is needed for the listing of critical rather than absolute maximum stresses in the 

result summary for piles consisting of materials of different strengths (for example, 

for a concrete pile with steel stinger, the concrete section may be more important to 

investigate for stresses then the steel stinger, even though the steel stinger may 

have stresses which are 10 times higher than the concrete stresses). Pile sections 

for which the Critical Index is set 1 are included in the search for a maximum tension 

and compression stresses for final output.

Figure 12-42 Data entry screen for non-uniform piles. 

Several infrequently used options concerning primarily the pile model can be 

accessed in Options/General Options/Numeric (see Figure 12-43). An important but 

somewhat different pile option leading to an alternate type of analysis is the Residual 

Stress Analysis (RSA). The input number indicates the maximum number of repeat 

analyses (or blows) allowed, with a “1” representing the absolute limit of 100 cycles. 

Potentially important for large piles is the input of an adjusted Hammer and Pile 

Gravity. The default gravitational constant is 32.17 ft/s2

. The default value would 

represent a vertically driven pile above the water table. If the static weight is less 
242

due to either pile inclination or buoyancy, this value should be reduced using 

Options/General Options/Numeric. Note that the hammer and pile mass magnitudes 

will not be affected by this change of their gravitational constant.

Soil parameter options, like pile options, allow for increased input flexibility. Under 

Options/Soil Parameters, it is possible to enter individual nominal resistance values, 

damping, and quake values for each pile segment. Use of these options causes the 

corresponding field labels in the soil input section of the Main Input Form to read 

“Variable.” 

Figure 12-43 Numeric options window. 

The Damping Option, accessible in Options/General Options/Damping (Figure 

12-44), is rarely used for routine applications and is more useful for the researcher. 

In most instances, the Soil Damping is set to Smith damping. If the Residual Stress 

Analysis is performed or a vibratory hammer is analyzed then the Smith viscous 

damping option should be selected. Also Refined Wave Equation analyses 

sometimes have to done with Smith viscous damping for a good match. Hammer 

Damping and Pile Damping Options have been preset to a percentage of the 

impedance of the ram and hammer cushion and the pile, respectively, though the 

preset values may be replaced with small non-negative integers. Given a negative 

input, the program will read a zero value. For the pile material, the program 

automatically chooses values of 1, 2 and 5 for steel, concrete and wood, 

respectively. Another pile material (e.g., plastic piles) may require other, possibly 
243

higher inputs. While not enough is known about these parameters, their effect on 

the computed results is relatively insignificant.

Figure 12-44 Damping options window. 

The Stroke Option, important for diesel hammers, is accessible in Options/General 

Options/Stroke (see Figure 12-45). For any diesel hammer, the stroke is a function 

of fuel settings, pile mass and stiffness, and soil resistance. The stroke option 

allows the user to control whether the program will analyze a fixed stroke or 

calculate the stroke (default) based on the combustion pressure provided in either 

the hammer database or the user modified input. A fixed stroke can either be 

analyzed with an iteratively adjusted combustion pressure, such that upstroke 

equals down stroke, or with a single impact whose upstroke is then potentially 

different from its down stroke. On the Main Input Form below the Hammer 

Information window, this selected stroke option is identified. The selection of the 

Stroke Option is particularly important for Inspector’s Chart analyses, and the 

reading of the associated Help is strongly recommended.
244

Figure 12-45 Stroke options window for diesel hammers. 

The Stroke Options window also allows for a selection of Fuel Settings for those 

diesel hammers that have stepwise adjustable fuel pumps. Alternatively, the 

corresponding fractions of the maximum combustion pressures can be selected on 

the Main Input Form. As noted in Section 12.6.3, analyzing a hammer with a high 

combustion pressure, even if the high stroke is the result of pre-ignition, may lead to 

high calculated transferred energies and, therefore, non-conservative nominal 

resistance predictions and conservative stress predictions. On the other hand, if the 

observed hammer stroke is relatively low and friction (which should be modeled with 

a reduced hammer efficiency) has been eliminated as a reason for the low stroke, a 

reduced combustion pressure presents a reasonable analysis option. Because of 

the potential for an overprediction of nominal resistance due to excessive pressure 

adjustment, the Inspector’s Chart option does not increase the combustion pressure 

above the value in the hammer data file despite the presence of any values for high 

stroke analyses in the ”Convergence of pressure with fixed stroke.”
245

12.8 WAVE EQUATION ANALYSIS OUTPUT AND INTERPRETATION

The GRLWEAP program offers several output options that may be invoked or 

modified using Options, General Options, Output as displayed in Figure 12-46. The 

box labeled Type allows for selection of certain variables (e.g., force, velocity, 

stress) at a number of different segments. The user may opt to plot these variables 

as a function of time or create a table for transfer to other programs. Of particular 

interest is the plotting of pile top force and proportional velocity vs. time for 

comparison with dynamic measurements.

The Numerical box underneath allows for the control of the numerical output in one 

of three means. Choosing Minimum (default for drivability analyses) will exclude the 

extrema tables that are included in the Normal output selection. The extrema tables 

are very helpful when investigating the location of maximum stress values, and even 

though they may make the output very long, it is often desirable to revert to the 

Normal option, even for drivability analyses. Another worthy candidate for the 

normal output is the multi-material pile. The final numerical option, Debug generates 

so much numerical output that it is rarely needed for real applications.

After an analysis has been run, clicking the O (Output) icon transfers control to the 

output program in which several output modes (depending on output and analysis 

options) of the Project Summary (containing several important parameters and title 

components) become available: Bearing Graph or Drivability, Variables vs. Time

and Numeric results. 
246

Figure 12-46 Output options window. 

The Numerical GRLWEAP Output, or Numeric results, is the most important output. 

It begins with a listing of file names used for input and the input file (*.GWW). There 

follows a disclaimer pointing out some of the uncertainties associated with wave 

equation analyses. The user is urged to check that the correct data file is used and 

consider the disclaimer when drawing conclusions from analysis results.

The first page of output, shown in Figure 12-47, lists the hammer and drive system 

components used in the analysis. Hence hammer model, stroke, and efficiency, 

helmet weight, and hammer and pile cushion properties including thickness, area, 

elastic modulus, and coefficient of restitution are but a few of the input details printed 

on this page.
247

Figure 12-47 Hammer model, hammer options, and driving system output. 

The second page of output, presented in Figure 12-48, summarizes the pile and soil 

model used in the analysis. A brief summary of the pile profile is provided at the top 

of the page and includes the pile length, area, modulus of elasticity, specific weight, 

perimeter, material strength (normally 0 for uniform piles), wave speed, and pile 

impedance (EA/C). A detailed summary of the pile and soil model follows the pile 

profile. The detailed pile model includes the number of pile segments, their weight 

and stiffness, and any compression slacks (C-Slk) or tension slacks (T-Slk) with 

associated coefficient of restitution (C.O.R.). The listing also shows segment bottom 

depth (LbTop) and the averages of both segment circumference and cross sectional 

area. The summarized soil model includes the soil static soil resistance distribution 

(Soil-S), the soil damping parameters (Soil-D) along the shaft and at the pile toe, as 
248

well as the soil quakes along the shaft and at the pile toe. Additional pile and soil 

modeling options, including the percent shaft resistance, are summarized below in 

Figure 12-48. 

On the third page, shown in Figure 12-49, an extrema table is printed for each pile 

segment number. This extrema output (Table 12-10) is printed for each analyzed 

nominal resistance and includes:

Table 12-10 GRLWEAP Extrema Output

Abbreviation Description

No Pile segment number

mxTForce Maximum tension (negative) force in kips

mxCForce Maximum compression force in kips

mxTStrss Maximum tension (negative) stress in ksi

mxCStrss Maximum compression stress in ksi

max V Maximum velocity in ft/s

max D Maximum displacement in inches

max Et Maximum transferred energy in kip-ft

The "t" values following the extreme values are times in milliseconds relative to 

hammer impact. Note that tension is shown as a negative in these tables. For the 

analysis of diesel hammers, the iteration on hammer stroke is indicated beneath the 

extrema table information followed by the maximum combustion pressure analyzed 

in psi.

For bearing graph analyses, GRLWEAP concludes by printing a summary table for 

all input nominal resistances (ultimate capacities) after the extrema table listing. The 

summary table is illustrated in Figure 12-50 and includes the analyzed nominal 

resistance (ultimate capacity, Rut) and corresponding penetration resistance (Bl Ct) 

for blow count, analyzed hammer stroke (for diesel hammers, both the down stroke 

and the rebound stroke), maximum tension stress (negative, Ten Str), maximum 

compression stress (Comp Str), maximum transferred energy (ENTHRU), and, for 

diesel hammers, hammer operating speed (Bl Rt). The indicators "i t" locate where 

(pile segment number) and when (time after impact in m/s) the extreme stress 

values occur.
249

Figure 12-48 Pile, soil, and analysis options. 
250

Figure 12-49 Extrema table output. 

Figure 12-50 Final summary for bearing graph analysis. 

Review of the "printed output" can be accomplished on the computer screen before 

printing. This review is extremely important as it can point out inadvertent omissions 

or erroneous input data. The reviewer should carefully check ram weight, stroke, 

efficiency, cushion stiffness, pile mass and stiffness values, soil parameters, etc. 

Furthermore, any error messages or warnings issued by the program should be 

checked for relevance to the results.
251

12.9 PLOTTING OF WAVE EQUATION RESULTS

The summary table results are usually presented in the form of a bearing graph 

relating the nominal resistance (ultimate capacity) to the pile penetration resistance. 

Plotting can be done by program built-in plotting routines or by saving the copying 

Numerical Output data to the clipboard and then pasting it in some other plotting 

program. The GRLWEAP program provides for the plotting of:

• one or two bearing graphs in the same plot, 

• an Inspector’s Chart,

• drivability results (e.g., blow count, nominal resistance, stresses, stroke, and 

transferred energy vs. depth) for one or two gain/loss factors, and 

• the plotting of selected variables vs. time (e.g., forces, velocities, displacements 

etc.) 

The wave equation bearing graph or inspector’s chart should be provided for the 

resident construction engineer, pile inspector, and the contractor.

12.10 SUGGESTIONS FOR PROBLEM SOLVING

Table 12-11 summarizes some of the field problems that can be solved through use 

of wave equation analysis. Field problems may arise due to soil, hammer, driving 

system, and pile conditions that are not as anticipated or unknown. Of course, all 

possibilities cannot be treated in this summary. Sometimes, the performance of the 

wave equation program may produce an unexpected or apparently useless result 

and a corrective action may be required. A number of such problems together with 

suggested solutions are listed in Table 12-12. Further information may also be 

found in PDI (2010) and in the program’s Help Menu.
252

Table 12-11 Suggested Use of the Wave Equation to Solve Field Problems

Problem Solution

Concrete pile spalling or 

slabbing near pile head.

Perform wave equation analysis; find pile head stress 

for observed blow count and compare with allowable 

stresses. If high calculated stress, add pile cushioning.

If low calculated stress, investigate pile quality, 

hammer performance, hammer-pile alignment.

Concrete piles develop 

complete horizontal cracks 

in easy driving.

Perform wave equation analysis; check tension 

stresses along pile (extrema tables) for observed blow 

counts. If high calculated tension stresses, add 

cushioning or reduce stroke. If low calculated tension 

stresses, check hammer performance and/or perform 

measurements.

Concrete piles develop 

complete horizontal cracks 

in hard driving.

Perform wave equation analysis; check tension 

stresses along pile (extrema table). If high calculated

tension stresses, consider heavier ram. If low 

calculated tension stresses, take measurements and 

determine quakes, which may be higher than 

anticipated.

Concrete piles develop 

partial horizontal cracks in 

easy driving.

Check hammer-pile alignment since bending may be 

the problem. If alignment appears to be normal, 

tension and bending combined may be too high; 

solution as for complete cracks.

Steel pile head deforms or 

timber pile top mushrooms.

Check helmet size/shape; check steel strength; check 

evenness of pile head, banding of timber pile head. If 

okay, perform wave equation and determine pile head 

stress. If calculated stress is high, reduce hammer 

energy (stroke) for low blow counts; for high blow 

counts different hammer or pile type may be required. 

Unexpectedly low blow 

counts during pile driving.

Investigate soil borings; if soil borings do not indicate 

soft layers, pile may be damaged below grade.

Perform wave equation and investigate both tension 

stresses along pile and compression stresses at toe. If 

calculated stresses are acceptable, investigate 

possibility of obstructions / uneven toe contact on hard 

layer or other reasons for pile toe damage.
253

Table 12-11 Suggested Use of the Wave Equation to Solve Field Problems 

(Continued)

Problem Solution

Higher blow count than 

expected.

Review wave equation analysis and check that all 

parameters were reasonably considered. Check 

hammer and driving system. If no obvious defects are 

found in driving system, field measurements should be 

taken. Problem could be pre-ignition, preadmission, 

low hammer efficiency, soft cushion, large quakes, 

high damping, greater soil strengths, or temporarily 

increased soil resistance with later relaxation (perform 

restrike tests to check).

Lower blow count than 

expected.

Probably soil resistance is lower than anticipated.

Perform wave equation and assess soil resistance.

Perform restrike testing (soil resistance may have 

been lost due to pile driving), establish setup factor 

and drive to lower nominal resistance. Hammer 

performance may also be better than anticipated, 

check by measurement.

Diesel hammer stroke 

(bounce chamber 

pressure) is higher than 

calculated.

The field observed stroke exceeds the wave equation 

calculated stroke by more than 10%. Check that 

hammer was set to correct setting. Compare 

calculated and observed blow counts. If observed are 

higher, soil resistance is probably higher than 

anticipated. If blow counts are comparable, reanalyze 

with higher combustion pressure to match observed 

stroke and assure that pre-ignition is not a problem, 

e.g., by measurements.

Diesel hammer stroke 

(bounce chamber 

pressure) is lower than 

calculated.

The field observed stroke is less than 90% of the 

stroke calculated by the wave equation. Check that 

hammer was set to correct setting. Check that ram 

friction is not a problem (ram surface should have well 

lubricated appearance). Compare calculated and 

observed blow count. If observed one is lower, soil 

resistance is probably lower than anticipated. If blow 

counts are comparable, reanalyze with lower 

combustion pressure to match observed hammer 

stroke.
254

Table 12-12 Wave Equation Analysis Problems

Problem Solution

Cannot find hammer in 

data file.

Contact the hammer manufacturer or the author(s) of 

the wave equation program. Pile Dynamics, Inc., for 

example, regularly updates and posts its hammer data 

file on its web page. Alternatively, the user may utilize 

a hammer of same type and of similar ram weight and 

energy rating and modify its data to match the unlisted 

hammer’s specifications as closely as possible.

Cannot find an acceptable 

hammer to drive pile within 

driving stress and 

penetration resistance 

limits.

Both calculated stresses and blow counts are too high.

Increase pile impedance or material strength or 

redesign for lower capacities. Alternatively, check 

whether soil has potential for setup. If soil is fine 

grained or known to exhibit setup gains after driving, 

then end of driving nominal resistance may be chosen 

lower than required. Nominal resistance should be 

confirmed by restrike testing or static load testing.

Diesel hammer analysis 

with low or zero transferred 

energies.

Probably soil resistance too low for hammer to run. Try 

higher capacities.

Unknown hammer energy 

setting.

Perform analyses until the cushion thickness/hammer 

energy setting combination yields acceptable stresses 

with minimum cushion thickness. Specify that the 

corresponding cushion thickness and hammer fuel 

setting be used in the field and their effectiveness 

verified by measurements.

Cannot find a suggested 

set of driving system data.

Contact contractor, equipment manufacturer, or use 

data for similar systems.

Unknown pile cushion 

thickness. 

Perform analyses until cushion thickness/hammer 

energy setting combination is found that yields 

acceptable stresses with minimum cushion thickness.

Specify that this thickness be used in the field and its 

effectiveness verified by measurements.

Calculated pile cushion 

thickness is uneconomical.

In order to limit stresses, an unusually thick pile 

cushion was needed for pile protection. Try to analyze 

with reduced energy settings. For tension stress 

problems, energy settings often can be increased after 

pile reaches sufficient soil resistance.
255

Table 12-12 Wave Equation Analysis Problems (Continued)

Problem Solution

Calculated driving times 

are unrealistically high or 

low.

The calculation of driving times is very sensitive, 

particularly at high blow counts. Use extreme caution 

when using these results for cost estimation. Also, no 

interruption times are included and the estimate is only 

applicable to non-refusal driving.

Wave equation calculated 

energy and/or forces are 

difficult to match with field 

measurements.

In general, it is often difficult to make all measured 

quantities agree with their calculated equivalents. A 

10% agreement should be sufficient. Parameters to 

be varied include hammer efficiency, diesel hammer 

combustion pressure, external combustion hammer 

stroke, coefficients of restitution, hammer cushion 

properties for steel piles, pile cushion properties for 

concrete piles, and pile top properties for timber piles. 

Resistance distribution also may affect the pile top 

forces.

12.11 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS

One of the primary advantages of wave equation analysis is the ability to evaluate 

the drivability and suitability of a pile foundation design early in the design process. 

This allows economic evaluations on pile section selection to be easily and rationally 

decided based on predicted driving stresses and penetration resistances. For the 

contractor, wave equation analysis provides one of the best methods for equipment 

selection, determining the need for pile installation aids, and selecting pile cushion 

thickness. For the engineer, wave equation analysis provides the best method for 

hammer approval. No other computational tool for the simulation of the dynamic pile 

driving event is able to match the convenience and realism of this approach. Once 

the necessary information for program input is in hand, (closest boring, proposed 

driving system, and foundation details), a wave equation analysis can be performed 

relatively quickly.

A disadvantage of wave equation analysis is the information needed on the soil 

conditions, pile and foundation details, and hammer system for the analyst to make 

informed program input selections. A dynamic formula is simpler and quicker, albeit 

also less accurate, as the only input information needed for formula use is generally 

the hammer energy. The program user should have knowledge of the software 
256

program, soil mechanics, and pile design and construction to understand and 

correctly apply results in unique and specific applications.

Wave equation limitations include, in some soil conditions, conservative estimates of 

the nominal resistance. Open end pipe piles or H-piles in granular profiles often 

behave differently under dynamic and static loading conditions. Under dynamic 

conditions, a plug may slip and produce additional shaft resistance. However under 

static loading, a plug may form over the full cross section resulting in soil resistance 

developing on a larger toe area. The wave equation model in this scenario may 

suitably predict the static behavior. The nominal resistance determined from wave 

equation analysis can be unconservative at low blow counts (less than 30 blows per 

foot) and overly conservative at very high blow counts (greater than 120 blows per 

foot).

12.12 PRACTICAL ISSUES AND CONSIDERATIONS

When confronted with the task of performing drivability analyses for a major bridge 

project, the question is often how many different analyses have to be performed to 

adequately represent the various pile, nominal resistance and soil conditions. 

Performing too many different analyses is not only time consuming it also may lead 

to confusions.

Nominal resistance determination by wave equation analysis or dynamic testing in 

soils with unknown long term behavior requires restrike testing. However, if a 

restrike test is performed, then it would add little cost to do dynamic testing at the 

same time and take advantage of a higher resistance factor. Thus wave equation 

based nominal resistance determination is only practical for small jobs where 

experience with similar conditions exist so that no restrike testing is necessary. In 

that case the use of the Inspector’s Chart is a valuable tool. For open end diesel 

hammers, this would, however, require that the hammer stroke is accurately 

monitored. Also in that case, care should be taken to check whether or not there are 

unusual changes in stroke and/or blow count vs. depth behavior which could, for 

example, be indications of an overheating hammer.

When analyzing the driving long concrete piles with a large number of hammer 

blows, it is important to consider that pile cushion replacement should be done after 

approximately 1500 hammer blows. A drivability analysis has to properly reflect the 

pile cushion replacement by a change of stiffness. It is not advisable to schedule 

such a cushion change shortly before the end of driving where a new cushion would 

cause a low energy transfer and therefore an inflated blow count.
257

When specifying a blow count criterion two considerations are important: if the end 

of penetration resistance is low because of a large hammer, check to see if the 

hammer energy can be sufficiently reduced to provide a driving criterion above 30 

blows/ft. Considering soil setup and allowing for a criterion with reduced nominal 

resistance may lead to a situation where the pile driving hammer is incapable to 

producing a blow count less than 120 blows/ft during a restrike. If nominal 

resistance verification by dynamic methods is important, then a larger hammer (e.g., 

a drop hammer) should be provided for the restrike test. Alternatively, the pile 

driving hammer has to be chosen large enough to produce restrike blow counts 

below 120 blows/ft.

When driving to a hard rock surface it would be unwise to specify the required 

penetration resistance in blows/foot. Instead it is recommended to specify a 

maximum pile penetration under a certain number of blows, for example, not more 

than 1⁄2 inch under 10 hammer blows. Bearing graphs and Inspector’s Charts can be 

provided with this alternative penetration resistance measurement.
258

REFERENCES

American Association of State Highway and Transportation Officials (AASHTO). 

(2014). AASHTO LRFD Bridge Design Specifications, US Customary Units, 

Seventh Edition, with 2015 Interim Revisions. American Association of State 

Highway and Transportation Officials, Washington, D.C., 1960 p.

Bielefeld, M.W. and Middendorp, P. (1992). Improved Pile Driving Prediction for 

Impact and Vibratory Hammers. Proceedings of the Fourth International 

Conference on the Application of Stresswave Theory to Piles, Balkema 

Publishers, A.A. The Hague, The Netherlands, pp. 395-399.

Blendy, M.M. (1979). Rational Approach to Pile Foundations. Symposium on Deep 

Foundations, ASCE National Convention.

Brown, D.A., and Thompson, W.R. (2015). Design and Load Testing of Large 

Diameter Open-Ended Driven Piles. A Synthesis of Highway Practice. 

NCHRP Synthesis 478, Transportation Research Board, Washington, D.C., 

2015.

Cheney, R.S. and Chassie, R.G. (2000). Soils and Foundations Workshop 

Reference Manual. FHWA HI-00-045, U.S. Department of Transportation, 

National Highway Institute, Federal Highway Administration, Washington, 

D.C., 358 p.

Goble, G.G. and Rausche, F. (1976). Wave Equation Analysis of Pile Driving – 

WEAP Program, FHWA IP-76-14.3., U.S. Department of Transportation, 

Federal Highway Administration, Office of Research and Development, 

Washington, D.C., Volumes I-IV.

Goble, G.G. and Rausche, F. (1986). Wave Equation Analysis of Pile Driving - 

WEAP86 Program. U.S. Department of Transportation, Federal Highway 

Administration, Implementation Division, McLean, VA, Volumes I-IV.
259

Hirsch, T.J., Carr, L. and Lowery, L.L. (1976). Pile Driving Analysis, TTI Program, 

IP-76-13. U.S. Department of Transportation, Federal Highway 

Administration, Offices of Research and Development, Washington, D.C., 

308 p.

Holloway, D.M. and Beddard, D.L. (1995). Dynamic Testing Results Indicator Pile 

Test Program – I-880. Proceedings of the 20th Annual Members Conference 

of the Deep Foundations Institute. Charleston, SC, pp. 105-126.

Middendorp, P. (2004). Thirty Years of Experience with the Wave Equation Solution 

Based on the Method of Characteristics. Proceedings of the Seventh 

International Conference on the Application of Stresswave Theory to Piles, 

Selangor, Kuala Lumpur, Malaysia, pp. 95-106.

PDI, Pile Dynamics, Inc. (2010). GRLWEAP Background Report, Version 2010.

Cleveland, OH, 149 p.

Rausche, F., Liang, L., Allin, R., and Rancman, D. (2004). Applications and 

Correlations of the Wave Equation Analysis Program GRLWEAP. 

Proceedings of the Seventh International Conference on the Application of 

Stresswave Theory to Piles, Selangor, Kuala Lumpur, Malaysia, pp. 107-123.

Rausche, F., Nagy, M., Webster, S., and Liang, I.L. (2009). CAPWAP and Refined 

Wave Equation Analyses for Drivability Predictions and Capacity Assessment 

of Offshore Pile Installations. Proceedings of the ASME Twenty Eighth 

International Conference on Ocean and Arctic Engineering, Paper No. 

OMAE2009-80163, Honolulu, HI, pp. 375-383.

Reiding, F.J., Middendorp, P., Schoenmaker, R.P., Middendorp, F.M., and Bielefeld, 

M.W. (1988). The FPDS-2, A New Generation of Foundation Diagnostic 

Equipment. Proceedings of the Third International Conference on the 

Application of Stress Wave Theory to Piles, Ottawa, Canada, BiTech 

Publishers, pp. 123-134.

Smith, E.A.L. (1960). Pile Driving Analysis by the Wave Equation. American 

Society of Civil Engineers (ASCE), Journal of the Soil Mechanics and 

Foundations Division, Vol. 86, No. 4, pp. 35-61.

Soares, M., de Mello, J., and de Matos, S. (1984). Pile Drivability Studies, Pile 

Driving Measurements. Proceedings of the Second International Conference 
260

on the Application of Stress-Wave Theory to Piles, Balkema Publishers, A.A., 

Stockholm, Sweden, pp. 64-71.

U.S. Army Corps of Engineers, (1991). Design of Pile Foundations, Engineering 

Manual EM 1110-2-2906. U.S. Department of the Army, Washington, D.C., 

184 p.
261

CHAPTER 13

DYNAMIC FORMULAS

13.1 BACKGROUND 

Engineers have attempted to find rational methods to determine the geotechnical 

resistance of a driven pile as long as they have been used for structure foundations. 

Initially, prediction methods were proposed using pile penetration observations made 

during driving. However, the only realistic measurement that could be obtained 

during driving was the pile set per blow. Thus energy concepts equating the 

potential energy of the hammer to the penetration resistance of the pile (set per 

blow) as it is driven through the soil were developed to estimate the geotechnical 

capacity or nominal pile resistance. In equation form this can be expressed as: 

bn = sRWh Eq. 13-1 

Where: 

W = ram weight.

h = ram stroke.

Rn = nominal resistance.

sb = set per blow.

These types of expressions are known as dynamic formulas. Because of their 

simplicity, dynamic formulas have been widely used for many years. Numerous 

dynamic formulas have also been proposed over time and some include 

consideration of pile weight, energy losses in drive system components, pile 

temporary compression, and other factors. Whether simple or more complex 

dynamic formulas are used, the nominal resistances determined by dynamic 

formulas have generally shown poor correlations and wide scatter when statistically 

compared with the nominal resistances determined by static load test results. 

AASHTO (2014) LRFD Bridge Design Specifications include two dynamic formulas; 

the FHWA modified Gates formula, discussed herein in Section 13.3.1, and an 

AASHTO modified version of the Engineering News formula, discussed further in 

Section 13.3.2. Both the AASHTO design and construction specifications state a 

dynamic formula should not be used when the nominal resistance exceeds 600 kips.
262

13.1.1 Historical Accuracy of Dynamic Formulas

Wellington proposed the popular Engineering News formula in 1892. It was 

developed for evaluating the nominal resistance or capacity of timber piles driven 

primarily by drop hammers in sand. Concrete and steel piles were unknown at that 

time, as were many of the pile hammer types and hammer sizes used today. 

Therefore, it should be of little surprise that the formula performs poorly in predicting 

the capacity or nominal resistance of the pile foundations used today.

The inadequacies of dynamic formulas have been known for a long time. In 1941, 

an ASCE committee on pile foundations assembled the results of numerous static 

load tests along with the predicted capacities from several dynamic formulas, 

including the Engineering News, Hiley, and Pacific Coast formulas. The mean 

failure load of the load test database was 91 tons. After reviewing the database, 

Peck (1942) proposed that a new and simple dynamic formula could be used that 

stated the capacity of every pile was 91 tons. Peck concluded that the use of this 

new formula would result in a prediction statistically closer to the actual pile capacity 

than that obtained by using any of the dynamic formulas contained in the 1941 

study. A more detailed discussion of both the 1941 ASCE debate as well as the 

inadequacies of dynamic formulas can be found in Likins et al. (2012).

Chellis (1961) noted that the actual factor of safety obtained by using the 

Engineering News formula varied from as low as 1⁄2 to as high as 16. Sowers (1979) 

reported that the safety factor from the Engineering News formula varied from as low 

as 2/3 to as high as 20. Fragasny et al. (1988) in the Washington State DOT study 

entitled "Comparison of Methods for Estimating Pile Capacity" found that the Hiley, 

Gates, Janbu, and Pacific Coast Uniform Building code formulas all provide 

relatively more dependable results than the Engineering News formula.

As part of a FHWA research project, Rausche et al. (1996) compiled a database of 

static load test piles that included pile capacity predictions using the FHWA 

recommended static analysis methods, preconstruction and refined wave equations, 

as well as dynamic measurements coupled with CAPWAP signal matching analysis. 

The reliability of the various capacity prediction methods were then compared with 

the results of the static loading tests. The results of these comparisons are 

presented in Figure 13-1 in the form of probability density function curves versus the 

ratio of predicted load over the static load test result. The mean values and 

coefficients of variation for the methods studied are presented in Table 13-1. The 

closer the mean value of the ratio of the predicted/static load test result is to 1.0 and 

the smaller the coefficient of variation (COV) the more reliable the method. 
263

Prediction method performance using driving observations of blow count and 

hammer stroke are identified as EOD for end of driving observations or BOR for 

beginning of restrike.

In the 1998 version of the FHWA pile manual, the database compiled by Rausche et 

al. (1996) was modified to include resistance predictions from the allowable load 

version of the Engineering News as well as the FHWA Modified Gates dynamic 

formulas at both the end of driving and beginning of restrike. The database for the 

dynamic formulas was also expanded and included additional data sets. The 

allowable load determined using the Engineering News formula in this study was 

compared to one half of the nominal resistance determined from the static load test, 

while the nominal resistance from the Modified Gates formula was compared directly 

to the nominal resistance determined from the static load test. The correlation 

results of the dynamic formulas are included in Table 13-1. 

Based on the end of driving data, the Engineering News formula had a mean value 

of 1.22 and a coefficient of variation of 0.74, while the Modified Gates had a mean 

value of 0.96 with a coefficient of variation of 0.41. The coefficient of variation is the 

standard deviation divided by the mean value. Hence, the greater a method's mean 

value is from 1.0 the lower the accuracy of the method, and the larger the coefficient 

of variation the less reliable the method. Table 13-1 clearly shows the Engineering 

News formula has a tendency to overpredict capacity. The higher coefficient of 

variation also suggests that the Engineering News formula is significantly less 

reliable than the Modified Gates formula.

Table 13-1 also illustrates that evaluation of pile capacity, by either Gates or 

Engineering News dynamic formula from restrike set and energy observations, has a 

significant tendency to overpredict capacity. The Engineering News formula 

capacity results, from restrike observations, had a mean value of 1.89 and a 

coefficient of variation of 0.46. The Modified Gates formula capacity results, from 

restrike observations, had a mean value of 1.33 and a coefficient of variations of 

0.48. 

If the static load test failure loads are divided by the Engineering News allowable 

design loads, the database indicates an average factor of safety of 2.3 as compared 

to the factor of safety of 6.0 theoretically included in the allowable load version of the 

formula. More important, the actual factor of safety from the Engineering News 

formula ranged from 0.6 to 13.1. This lack of reliability causes the Engineering 

News formula to be ineffective as a tool for estimating capacity. The fact that 12% of 

the database has a factor of safety of 1.0 or less is also significant. However, 
264

complete failure of a bridge due to inadequate geotechnical resistance determined 

by Engineering News formula is unusual. The problem usually is indicated by long 

term damaging settlements which occur after construction.

Figure 13-1 Log normal probability density function for four resistance predictions 

(after Rausche et al. 1996).

Dynamic formulas were historically used on small projects where conservative 

design load estimates, greater foundation redundancy, and resultant reserve 

foundation capacity helped mitigate some foundation performance problems. This 

hidden reserve resistance has been largely reduced in LRFD designs that utilize 

fewer, larger size piles, with higher nominal resistances.

The version of the Engineering News formula in the AASHTO (2014) LRFD Bridge 

Design Specifications is a modified from the historical allowable load version and 

calculates an ultimate capacity or nominal resistance. However, inherent problems 

with dynamic formulas remain as discussed in Section 13.1.2.
265

Table 13-1 Mean Values and Coefficients of Variation for Various Methods

Prediction Method Status Mean C.O.V. # Piles

Standard WEAP* BOR 1.22 0.35 99

Hammer Performance Adjusted WEAP* BOR 1.16 0.35 99

CAPWAP* BOR 0.92 0.22 99

Static Analysis* - 1.30 0.68 89

Engineering News Formula EOD 1.22 0.74 139

Engineering News Formula BOR 1.89 0.46 122

Modified Gates Formula EOD 0.96 0.41 139

Modified Gates Formula BOR 1.33 0.48 122

* From Rausche et al. (1996)

EOD = End of Driving, BOR = Beginning of Restrike 

13.1.2 Basic Limitations with Dynamic Formulas

Dynamic formulas have limitations, and are therefore less reliable than other field 

methods for nominal resistance verification. The basic limitations associated with 

pile driving formulas can be traced to the modeling of each component within the pile 

driving process: the driving system, the pile, and the soil. Dynamic formulas poorly 

represent the driving system and the energy losses of drive system components. 

Dynamic formulas also assume a rigid pile, thus neglecting pile axial stiffness effects 

on drivability, and further assume that the soil resistance is constant and 

instantaneous to the impact force. A more detailed discussion of these limitations is 

presented below.

The derivation of most formulas is not based on a realistic treatment of the driving 

system. Most formulas only consider the potential energy of the driving system. 

The variability of equipment performance is typically not considered. Driving 

systems include many elements in addition to the ram, such as the anvil for a diesel 

hammer, the helmet, the hammer cushion, and for a concrete pile, the pile cushion. 

These components affect the distribution of the hammer energy with time, both at 

and after impact, which influences the magnitude and duration of peak force. The 

peak force and its duration determine the ability of the driving system to advance the 

pile into the soil.
266

Dynamic formulas also assume that the pile is rigid and its length is not considered. 

This assumption completely neglects the pile's flexibility, which affects its ability to 

penetrate the soil. The energy delivered by the hammer sets up time-dependent 

stresses and displacements in the helmet, in the pile, and in the surrounding soil. In 

addition, the pile behaves, not as a concentrated mass, but as a long elastic rod in 

which stresses travel longitudinally as waves. Compression waves which travel to 

the pile toe are responsible for advancing the pile into the ground.

The soil resistance is also very crudely treated by assuming that it is a constant 

force. This assumption neglects the characteristics of real soil behavior. The 

dynamic soil resistance is the resistance of the soil to rapid pile penetration 

produced by a hammer blow. This resistance is in no way similar to the static soil 

resistance. However, most dynamic formulas consider the resistance during driving 

equal to the nominal resistance or capacity, and do not consider the dynamic 

behavior of the soil during pile penetration. The rapid penetration of the pile into the 

soil during driving is resisted not only by static friction and cohesion, but also by the 

soil viscosity, which is comparable to the viscous resistance of liquids against rapid 

displacement under an applied force. The net effect is that the driving process 

creates dynamic soil resistance forces along the pile shaft and at the pile toe, due to 

the high shear rate. The soil resistance during driving, from the combination of 

dynamic soil resistance and available static soil resistance, is generally not equal to 

the static soil resistance under static loads.

13.2 RESISTANCE FACTORS FOR DYNAMIC FORMULAS

Resistance factors applicable to the nominal axial geotechnical resistance 

determined from select dynamic formulas are provided in Table 13-2. The 

resistance factor varies depending on the dynamic formula and, in some cases, pile 

type. Only the FHWA modified Gates formula and AASHTO modified Engineering 

News formula have resistance factors in AASHTO. If a dynamic formula is used to 

establish driving criterion, AASHTO (2010) LRFD Bridge Construction Specifications 

Article 4.4.4.5 recommends use of the FHWA modified Gates formula. If a dynamic 

formula other than the FHWA modified Gates formula or AASHTO modified 

Engineering News formula is used, AASHTO specifications state that it should be 

calibrated based on measured static load test results to obtain an appropriate 

resistance factor.

The lower reliability of dynamic formulas is supported by the resistance factors for 

dynamic formulas contained in the AASHTO (2014) LRFD Bridge Design 
267

Specifications. The AASHTO resistance factor for nominal resistances determined 

by dynamic formula is 0.40 for the FHWA modified Gates formula and 0.10 for the 

AASHTO modified Engineering News Formula. For non-redundant foundations 

consisting of 4 piles or less, the resistance factor associated with a given dynamic 

formula should be further reduced by 20%.

Table 13-2 Resistance Factors for Dynamic Formulas 

(modified from AASHTO 2014)

Formula Resistance Determination 

Method

AASHTO (2014)

Resistance Factor

FHWA Modified 

Gates Formula, φdyn

Nominal resistance determined 

using FHWA Modified Gates 

Formula at end of driving 

condition only. See Section 

13.3.1.

0.40 

AASHTO Modified 

Engineering News 

Formula, φdyn

Nominal resistance determined 

using AASHTO Modified 

Engineering News Formula at 

end of driving condition only. 

See Section 13.3.2.

0.10

WSDOT Dynamic 

Formula, φdyn

Nominal resistance determined 

using WSDOT Dynamic 

Formula at end of driving 

condition only. See Section 

13.3.3.1.

Not in AASHTO

See discussion in text.

MnDOT Dynamic 

Formula, φdyn

Nominal resistance determined 

using MnDOT Dynamic 

Formula. See Section 13.3.3.2.

Not in AASHTO

See discussion in text.
268

13.3 DYNAMIC FORMULAS

13.3.1 FHWA Modified Gates Formula

For small projects where a dynamic formula is used, AASHTO states that the FHWA 

Modified Gates formula is preferable, since it correlates better with static load test 

results. The FHWA Modified Gates formula below has been revised to reflect the 

nominal resistance in kips and includes the 80 percent efficiency factor on the rated 

energy, Ed, recommended by Gates. The specified units below must be used.

100)10(log75.1 Rndr = Ed 10 Nb − Eq. 13-2 

Where: 

Rndr = nominal driving resistance (kips).

Ed = developed hammer energy (ft-lbs). If ram velocity is not measured, it 

may be assumed equal to the potential energy of the ram in the form 

of ram weight, W, (lbs) times stroke height, h (ft).

Nb = pile penetration resistance (blows/inch).

It is often desirable for construction inspection personnel or contractors to know the 

number of hammer blows per foot of pile penetration which will be required to obtain 

the specified nominal resistance. The FHWA modified Gates formula for this 

purpose can be re-written as follows: 

)10(12

x N ft = Eq. 13-3 

In which:

[( ) += 100 /( ]−1)75.1 Rx ndr Ed Eq. 13-4 

Where:

Nft = pile penetration resistance (blows/foot).

 x = exponent defined by Equation 13-4 and terms per Equation 13-2.

All dynamic formulas are empirically developed based the resistance predicted using 

end of drive blow count observations with the resistance determined from static load 

test results performed at a later time. The FHWA modified Gates formula therefore 

inherently includes time dependent resistance changes due to soil setup or 

relaxation. AASHTO (2014) recommends a resistance factor, φdyn, of 0.40 for the 

FHWA Modified Gates formula, and that the formula be used only for end of drive 

conditions.
269

13.3.2 AASHTO Modified Engineering News Formula

The Engineering News formula was developed to provide an estimated nominal 

resistance based upon hammer energy and the observed pile set. The formula uses 

the pile set during driving to empirically estimate the long term nominal resistance. 

Therefore, this dynamic formula also incorporates any time dependent soil 

resistance changes as a result of the empirical procedure. For this reason, restrike 

set observation should not be used to calculate the nominal resistance.

The AASHTO modified version for nominal resistance calculations based on end of 

driving conditions is presented in Equation 13-5. The specified units for energy and 

set must be used.

1.0

12

+ =

b

d

ndr

s

E

R Eq. 13-5 

Where:

Rndr = nominal driving resistance (kips).

Ed = developed hammer energy (ft-kips). 

sb = permanent pile set (inches).

As noted in Section 13.1.1, the Engineering News formula has long been recognized 

to be one of the least accurate and least consistent of the dynamic formulas. 

AASHTO specifications (2014) reflect this by assigning a resistance factor of 0.10 to 

the Engineering News formula and recommending that the formula be applied to 

only end of drive conditions.

13.3.3 Other Dynamic Formulas

A few state transportation agencies have developed their own dynamic formulas and 

associated resistance factors through calibration with static load test databases. 

Hammer type, pile type and size, as well as geologic conditions were considered in 

the dynamic formulas developed by the Washington State DOT and Minnesota DOT.

13.3.3.1Washington State DOT Pile Driving Formula

The Washington State DOT utilized a database of 141 static pile load results from 

Paikowsky et al. (2004) to evaluate nominal resistance predictions made by the 

FHWA Modified Gates formula relative to the static load test results. The original 

WSDOT intent was to slightly modify and improve the FHWA modified Gates 
270

formula. However, enough changes were made that a new dynamic formula was 

developed and named the WSDOT pile driving formula. The formula is presented in 

Equation 13-6 and the associated research study is presented in Allen (2005).

)10ln(6.6 ndr = eff d NEFR b Eq. 13-6 

Where:

Rndr = nominal pile resistance measured during driving (kips).

Feff = Hammer efficiency factor.

0.55 for air/steam hammers on all pile types.

0.35 for closed end diesel hammers on all pile type.

0.47 for open end diesel hammers on steel piles.

0.37 for open end diesel hammers on timber or concrete piles.

0.58 for hydraulic hammers on all pile types.

0.28 for drop hammers on all pile types.

Ed = developed energy, equal to W * h (ft-kips). 

W = ram weight (kips).

h = observed ram stroke (feet).

Nb = number of blows for 1.0 inch of pile permanent set, averaged over 

the last four inches of driving.

Table 13-3 summarizes the pile types, the range in pile section size, the range in 

load test failure load, and the number of data sets for a given pile type in the 

correlation database. Based on the database used to develop the formula, Allen 

estimated the average amount of resistance from soil setup included in the formula 

resistance prediction is about 30 to 70%. Hence, the formula should not be used for 

nominal resistance assessments using restrike observations since setup is already 

included in the dynamic formula correlation calibration.

Monte Carlo simulations were used for reliability analyses to estimate the reliability 

index, β, and the resistance factor needed to achieve a target reliability index value 

of 2.3 or 3.0, Allen (2005). Based on a target β of 2.3, WSDOT recommended a 

resistance factor, φdyn, of 0.55 be used with this formula for redundant foundations. 

For non-redundant foundations, defined as 4 piles or less, the WSDOT 

recommended resistance factor is 0.45 based on a target β of 3.0. The WSDOT 

dynamic formula, either as presented in Equation 13-6 or in a slightly modified 

version, has been adopted by the state transportation agencies in Washington, New 

Mexico, and Illinois.
271

Table 13-3 Pile Types and Sizes Contained in WSDOT Formula Database

Pile

Type

Range in

Pile Size

Range in Static 

Load Test Failure

Load for Pile Type 

(kips) 

Number of 

Datasets

Precast Concrete 12 to 36 inch 308 to 1797 49

Closed End Pipe 10 to 48 inch O.D. 237 to 1300 46

H-pile 10 to 14 inch 214 to 1239 29

Open End Pipe 24 to 60 inch O.D. 586 to 1984 9 

Concrete Cylinder 20 to 54 inch O.D. 324 to 1452 5 

Monotube N/A 227 to 463 2 

Timber N/A 200 1 

13.3.3.2 Minnesota Pile Formula 2012 (MPF12)

Paikowsky et al. (2009) combined pipe pile and H-pile load test information from 

Database PD/LT 2000, developed for the NCHRP Report 507 LRFD calibration 

study, with additional pipe and H-pile correlation data gathered from MnDOT 

practice. The original MnDOT/LT 2008 database included 40 H-pile, 65 closed end 

pipe pile, and 12 open end pipe pile data sets typical of the pile types, sizes, and soil 

conditions encountered in MnDOT practice. With time, additional datasets were 

added to the database. The expanded database was used for development and 

calibration of a new dynamic formula named the Minnesota Pile Formula 2012 

(MPF12).

In using the formula, it was recommended that the hammer energy (W)(h) be limited 

to 85% of the hammer manufacturer’s maximum rated energy in cases where the 

hammer energy exceeded the 85% value. The Minnesota Pile Formula 2012 

(MPF12) for use on steel and concrete piles is presented in Equation 13-7.

 Eq. 13-7 
272

Where:

Rndr = nominal driving resistance (kips).

 W = ram weight (lbs). 

h = observed ram stroke (feet).

sb = permanent set of last hammer blow (inches).

Wh = not to exceed 85% of manufacturer’s maximum rated energy (ft-lbs).

The first order second moment method (FOSM) procedure and Monte Carlo 

simulations were used for reliability analyses to estimate a target reliability index 

value of 2.33 for redundant foundations. Based on these procedures, Paikowsky et 

al. (2014), recommended resistance factors, φdyn, of 0.50 for pipe piles, 0.60 for Hpiles, 0.50 for non-voided, prestressed concrete piles up to 24 inch in size, and 0.80 

for voided prestressed concrete piles 20 inch to 54 inch in size. It was 

recommended that the MPF12 formula be further modified as part of a later research 

study for the large, voided, concrete piles. All of the resistance factors 

recommended in the study were also based on the pile penetration resistance falling 

between 2 and 15 blows per inch.

The MPF12 dynamic formula was also evaluated for timber piles. It was noted that a 

modifier of 0.5 should be applied to the formula for timber piles due to the increased 

energy loss. The hammer energy (W)(h) should be limited to 85% of the hammer 

manufacturer’s maximum rated energy in cases where the hammer energy exceeds 

the 85% value. The resulting modified form of MPF12 for timber piles is as follows:















 =

b

ndr

s

Wh R 10 log 1000

20 Eq. 13-8 

Where:

Rndr = nominal driving resistance (kips).

W = ram weight (lbs).

h = observed ram stroke (feet).

sb = permanent set of last hammer blow (inches).

Wh = not to exceed 85% of manufacturer’s maximum rated energy (ft-lbs). 

Paikowsky et al. (2014) recommended a resistance factor, φdyn, of 0.60 be used for 

the nominal resistance determined with Equation 13-8 for timber piles. This 

resistance factor was also based on a pile penetration resistance of between 2 and 

15 blows per inch.
273

The research study proposed that the MPF12 dynamic formulas could be used for 

either end of driving or restrike conditions with no change in the resistance factor or 

formula. This is somewhat unusual as most dynamic formulas are limited to only 

end of drive applications since they are empirically correlated to load test results and 

therefore inherently include time dependent changes in the nominal geotechnical 

resistance. 

13.4 DYNAMIC FORMULA CASE HISTORY

To illustrate the variable performance of dynamic formulas compared to more 

reliable methods, a case history will be briefly discussed. A 50 feet long, 24 inch 

square, prestressed concrete pile was driven 45 feet below grade with an ICE I-46 

single acting diesel hammer. The factored load to be supported by the pile was 380 

kips. Soil conditions consisted of 30 feet of loose to medium dense sand overlying a 

5 feet thick layer of medium dense cemented sand and limestone. The cemented 

sand was underlain by the intended bearing layer of medium dense to dense sand. 

At the end of driving the test pile had a penetration resistance of 49 blows per foot at 

an average hammer stroke of 8.14 feet. The test pile was restruck 5 days after initial 

driving and had restrike penetration resistances of 4, 2, 2, and 2 blows per inch. The 

corresponding average hammer stroke at the beginning and end of this restrike was 

8.05 and 8.55 feet, respectively. An axial compression load test was performed on 

this pile one week after initial driving (3 days after the restrike). Following the static 

load test, the pile was again restruck. The penetration resistances per inch of the 

second restrike were 4, 3, 3, 3, 3, 3, 3, and 4 blows per inch. The average hammer 

stroke at the beginning of this second restrike was 7.21 feet.

Nominal resistance estimates from dynamic formulas as well as from wave equation 

analysis, dynamic testing with signal matching, and static load testing are 

summarized in Table 13-4. The wave equation nominal resistance predictions were 

obtained from a fixed stroke bearing graph analysis with the soil model determined 

using default values from the GRLWEAP ST soil model. Dynamic testing signal 

matching results are based on CAPWAP results. The static load test failure load 

was determined using the Davisson Offset limit criteria. 

For the presented case, the MnDOT dynamic formula provided the closest nominal 

resistance prediction to the static load test result based on inputting end of initial 

driving penetration resistance and stroke values into the four dynamic formulas. 

Overall the closest correlation to the static load test determined nominal resistance 
274

was achieved from wave equation analysis (1% underprediction) and dynamic 

testing with signal matching (4% underprediction). A comparison of the maximum 

factored resistance that could be supported based on the results from a given 

method is presented in Table 13-5. The factored dynamic test results with signal 

matching on beginning of restrike data correlated best to the maximum factored 

resistance determined from the static load test.
275

Table 13-4 Case History – Comparison of Calculated Nominal Resistances 

Nominal

Resistance

Method

Test

Condition

Resistance 

Factor

for Test 

Method

φdyn

Nominal

Resistance

(kips)

Difference 

Relative to 

Static Load 

Test 

Resistance

(%)

AASHTO EN Formula EOD 0.10 2872 + 463

AASHTO EN Formula BOR-1 - - - 2799 (2) + 449

AASHTO EN Formula EOR-1 - - - 1734 (2) + 240

AASHTO EN Formula BOR-2 - - - 2507 (2) + 391

FHWA Modified Gates Formula EOD 0.40 710 + 39

FHWA Modified Gates Formula BOR-1 - - - 701 (2) + 37

FHWA Modified Gates Formula EOR-1 - - - 570 (2) + 12

FHWA Modified Gates Formula BOR-2 - - - 658 (2) + 29

WSDOT Formula EOD 0.55 (1) 950 + 86

WSDOT Formula BOR-1 - - - 934 (2) + 83

WSDOT Formula EOR-1 - - - 806 (2) + 58

WSDOT Formula BOR-2 - - - 837 (2) + 64

MNDOT Formula EOD 0.50 (1) 585 + 15

MNDOT Formula BOR-1 0.50 (1) 579 + 14

MNDOT Formula EOR-1 0.50 (1) 485 - 5 

MNDOT Formula BOR-2 0.50 (1) 548 + 7

Wave Equation Analysis EOD 0.50 560 + 10

Wave Equation Analysis BOR-1 0.50 560 + 10

Wave Equation Analysis EOR-1 0.50 440 - 14

Wave Equation Analysis BOR-2 0.50 505 - 1 

Dynamic Test & Signal Matching EOD 0.65 600 + 18

Dynamic Test & Signal Matching BOR-1 0.65 523 + 3 

Dynamic Test & Signal Matching EOR-1 0.65 460 - 10

Dynamic Test & Signal Matching BOR-2 0.65 490 - 4 

Static Load Test - - - - - - 510 - - - 

 Notes: (1) – Resistance factor not in AASHTO. 

(2) – Method not recommended to be used in restrike condition by 

AASHTO or formula developer. 
276

Table 13-5 Case History – Comparison of Factored Resistance

Nominal

Resistance

Method

Test

Condition

Resistance 

Factor

for Test 

Method

φdyn

Maximum 

Factored 

Resistance

From 

Method

(kips)

Difference 

Relative to 

Load Test 

Determined 

Factored 

Resistance

(%)

AASHTO EN Formula EOD 0.10 287 - 13

AASHTO EN Formula BOR-1 - - - 280 (2) - 16

AASHTO EN Formula EOR-1 - - - 173 (2) - 48

AASHTO EN Formula BOR-2 - - - 251 (2) - 24

FHWA Modified Gates Formula EOD 0.40 284 - 14

FHWA Modified Gates Formula BOR-1 - - - 280 (2) - 15

FHWA Modified Gates Formula EOR-1 - - - 228 (2) - 31

FHWA Modified Gates Formula BOR-2 - - - 263 (2) - 21

WSDOT Formula EOD 0.55 (1) 522 + 58

WSDOT Formula BOR-1 - - - 514 (2) + 55

WSDOT Formula EOR-1 - - - 443 (2) + 34

WSDOT Formula BOR-2 - - - 460 (2) + 39

MNDOT Formula EOD 0.50 (1) 293 - 12

MNDOT Formula BOR-1 0.50 (1) 289 - 13

MNDOT Formula EOR-1 0.50 (1) 242 - 27

MNDOT Formula BOR-2 0.50 (1) 274 - 17

Wave Equation Analysis EOD 0.50 280 - 16

Wave Equation Analysis BOR-1 0.50 280 - 16

Wave Equation Analysis EOR-1 0.50 220 - 34

Wave Equation Analysis BOR-2 0.50 253 - 24

Dynamic Test & Signal Matching EOD 0.65 390 + 18

Dynamic Test & Signal Matching BOR-1 0.65 340 + 3

Dynamic Test & Signal Matching EOR-1 0.65 299 - 10

Dynamic Test & Signal Matching BOR-2 0.65 319 - 4 

Static Load Test - - - 0.75 332 - - - 

 Notes: (1) – Resistance factor not in AASHTO. 

(2) – Method not recommended to be used in restrike condition by

AASHTO or formula developer. 
277

13.5 ADVANTAGES, DISADVANTAGES, AND LIMITATIONS

Dynamic formulas offer a method to quickly estimate nominal resistance during 

driving. Because of the simple inputs such as hammer energy and pile set, relatively 

little engineering judgement is needed to perform the calculations. The primary 

advantages of dynamic formula use are the immediate availability of the driving 

criterion and the minimal delay to pile driving operations.

Conversely, dynamic formulas have several disadvantages. They do not consider 

the entire driving system (i.e. hammer, pile, and soil), variation in hammer 

performance, nor energy losses due to pile stiffness. However, the primary 

disadvantages are formula accuracy and reliability. Most shortcomings of dynamic 

formulas can be overcome by a more realistic analysis of the pile driving process, 

such a wave equation analysis. Dynamic testing and analysis is another tool which 

is superior to use of dynamic formulas.

AASHTO limits dynamic formula use to piles with a nominal resistance of 600 kips or 

less. This nominal resistance value is well beyond the typical nominal resistance 

values in historical databases evaluating dynamic formula performance. Other 

codes such as the International Building Code limit formula use to 160 kips because 

of their limitations. Dynamic formulas are based on Newtonian impact theory which 

is invalidated by the use of hammer and pile cushions.

13.6 PRACTICAL ISSUES AND CONSIDERATIONS

Use of a dynamic formula is often “justified” because of the small number of piles on 

a project or the presence of a consistent and hard bearing layer such as bedrock. 

On small projects with a limited number of piles, the cost of the extra pile length 

resulting from use of a less reliable nominal resistance verification method may be 

more economical than the testing cost or test method impact on the construction 

schedule. Similarly when piles are driven to a hard bearing layer, little additional pile 

length may be necessary to achieve the higher nominal resistance required by a 

dynamic formula and more reliable resistance verification methods may not be 

economically justified. However, driving stresses and their control must be 

considered in this situation and this assessment cannot be made using a dynamic 

formula.
278

The Indiana Department of Transportation evaluated installation costs on driven pile 

foundations for state bridge projects installed from 2009 to 2014. The pile 

foundations were installed using either the FHWA modified Gates dynamic formula 

or dynamic testing with signal matching. On the dynamic testing controlled projects, 

each test pile was dynamically monitored during both initial driving and during 

restrike with the restrike test conducted between 1 and 7 days after initial driving 

depending on the subsurface conditions. The cost per lineal foot of pile and the cost 

per kip of load support for pile foundation projects controlled by dynamic formula or 

dynamic testing with signal matching as reported by Zaheer et al. (2015) is 

summarized in Table 13-6. 

Table 13-6 INDOT Dynamic Formula and Dynamic Testing Comparison

Details 

FHWA 

Modified 

Gates 

Formula 

Dynamic 

Testing 

with Signal 

Matching 

Totals

Plan Contract Length (ft) 246,052 995,100 1,241,152

Paid Pile Length (ft) 216,664 937,873 1,154,517

Pile Length underrun or {overrun} (lf) 29,408 57,227 86,638

Pile Length underrun or {overrun} (%) 13.6% 6.1% 7.5%

Cost Paid to Contractor $11,653,634 $46,178,800 $57,832,434

Average Unit Cost (per lf) $ 53.79 $ 49.24 $ 50.09

Zaheer et al. (2015) determined that the use of a simple dynamic formula has 

several economic drawbacks. Their review also concluded the factored load carried 

by a dynamic formula controlled pile was 21% less than the factored load carried by 

a dynamically tested pile. For the same factored load, pile lengths determined 

through use of a dynamic formula were 10 to 20% greater than the length 

determined from dynamic testing. The cost per kip of supported structure load for a 

dynamic formula pile was 39% higher than that of a dynamically tested pile. Overall, 

the average unit cost per linear foot of dynamic formula installed pile was 9.2% more 

than the cost per foot of dynamic test with signal matching installed pile.
279

REFERENCES

Allen, T.M. (2005). Development of the WSDOT Pile Driving Formula and Its 

Calibration for Load and Resistance Factor Design (LRFD), WA-RD 610.1, 

Research Office, Washington State Department of Transportation, Olympia, 

WA, 45 p.

American Association of State Highway and Transportation Officials (AASHTO). 

(2014). AASHTO LRFD Bridge Design Specifications, US Customary Units, 

Seventh Edition, with 2015 Interim Revisions. American Association of State 

Highway and Transportation Officials, Washington, D.C., 1960 p.

Chellis R.D. (1961). Pile Foundations. Second Edition, McGraw-Hill Book Company, 

New York, NY, pp. 21-23.

Fragasny, R.J., Higgins, J.D. and Argo, D.E. (1988). Comparison of Methods for 

Estimating Pile Capacity, WA-RD 163.1, Washington State Department of 

Transportation, Olympia, WA, 62 p.

Paikowsky, S.G. (2004). with contributions from Birgisson, B., McVay, M., Nguyen, 

T., Kuo, C., Baecher, G., Ayyub, B., Stenersen, K., O.Malley, K., 

Chernauskas, L., and O’Neill, M., Load and Resistance Factor Design (LRFD) 

for Deep Foundations, NCHRP Report 507. Transportation Research Board, 

Washington, D.C., 76 p.

Paikowsky, S.G., Marchionda, C.M., O'Hearn, C.M., and Canniff, M.C. (2009). 

Developing a Resistance Factor for Mn/DOT's Pile Driving Formula, MN/RC 

2009-37. Minnesota Department of Transportation, St. Paul, MN, 294 p.

Paikowsky, S.G., Canniff, M., Robertson, S., and Budge, A.S. (2014). Load and 

Resistance Factor Design (LRFD) Pile Driving Project – Phase II Study, 

MN/RC 2014-16. Minnesota Department of Transportation, St. Paul, MN, 

514 p.

Peck, R.B. (1942). Discussion: Pile Driving Formulas. Proceedings of the American 

Society of Civil Engineers, Vol. 68, No. 2, pp. 905-909.
280

Rausche, F., Thendean, G., Abou-matar, H., Likins, G.E. and Goble, G.G. (1996). 

Determination of Pile Drivability and Capacity from Penetration Tests, 

DTFH61-91-C-00047, Final Report. U.S. Department of Transportation, 

Federal Highway Administration, McLean, VA, 432 p.

Sowers, G.F. (1979). Introductory Soil Mechanics and Foundations. Fourth Edition, 

Macmillan Publishing Co., Inc., New York, NY, pp. 531-533.

Wellington, A. (1892). Discussion of “The Iron Wharf at Fort Monroe, VA by J.B. 

Chucklee.” American Society of Civil Engineers (ASCE), Transactions, Vol. 

27, No. 543, pp. 129-172.

Zaheer, M., Salgado, R., Prezzi, M., and Han, F. (2015). INDOT/Purdue Pile Driving 

Method for Estimation of Axial Capacity. Presentation at 2015 Purdue Road 

School Transportation Conference and Expo.
281

CHAPTER 14

CONTRACT DOCUMENTS

14.1 OVERVIEW OF PLAN AND SPECIFICATION REQUIREMENTS

Pile foundations generally cannot be inspected after installation. Therefore, 

construction specifications and monitoring are of prime importance for a successful 

pile foundation. Preparation of the contract plans, plan details, and specifications 

related to piling issues are the responsibility of the foundation designer in 

cooperation with materials and construction personnel. Project plans should include:

• Location of piles.

• Pile numbering system to clearly identify each pile in group or bent.

• Pile type, section, and estimated length.

• Pile toe details, driving shoe, closure plate, etc.

• Pile splicing details.

• Pile cut off elevation.

• Pile cap connection details.

• Estimated pile toe elevation.

• Minimum pile toe elevation, if needed.

• Required pile batter and direction.

• Orientation of H-piles.

• Factored resistance, Rr. 

• Nominal resistance, Rn. 

• Nominal driving resistance, Rndr. 

• Location of subsurface borings.

• Results of subsurface exploration.

It is the designer's responsibility to confirm that plans and specifications have been 

prepared using compatible language. This is particularly true in defining the required 

nominal driving resistance, which is an important component of any driven pile 

specification. Problems can arise when plans provide only the factored resistance, 

the nominal resistance, or the nominal driving resistance without other details such 

as the resistance factor associated with the construction control method or 

anticipated losses in resistance due to scour, liquefaction, or relaxation. For 
282

example, plan statements such as "piles shall have a resistance of 300 kips," 

provide an unclear description of the contract requirements as the resistance could 

be interpreted as the factored resistance, nominal resistance, or nominal driving 

resistance. Plans should therefore clearly indicate the factored resistance, the 

resistance factor, the nominal resistance, the additional resistance from any 

unsuitable layers (scour, liquefaction and their elevations), resistance changes 

following driving (setup or relaxation effects) and the resulting nominal driving 

resistance.

This chapter includes a generic pile specification for highway projects that was 

originally developed with input from State and Federal bridge and geotechnical 

engineers and released in FHWA Geotechnical Guideline 13. It has been updated 

over time as necessary and modified from an ASD to LRFD specification. AASHTO 

LRFD Bridge Construction specifications (2010) provide a similar document with 

additional commentary. A good driven pile specification should include the basic 

components in Table 14-1. 

The intent of the attached generic specification is to provide highway designers and 

transportation agencies with a comprehensive driven pile specification. However, 

this specification is not intended to be used directly for project application without 

review and modification by the foundation designer and transportation agency. 

Commentary sections are included where appropriate to assist the foundation 

designer in tailoring the generic specification to project requirements. The 

commentary sections explain the reasons behind development of particular sections 

of the specification and the relationship of the specification requirements to 

necessary pile design or construction activities.

Note that only driven piles are addressed in the specification. Other deep foundation 

types such as drilled shafts require completely different construction controls and are 

therefore not appropriate for inclusion in this generic pile specification.

In conventional design-bid-build contracts, agency standard specifications are used. 

The agency warrants to the contractor that the drawings and prescriptive 

specifications are complete and free from error (agency takes the risk). In designbuild contracts, specifications are performance based which allows the designbuilder to use their design and construction expertise to satisfy project requirements. 

The design-builder warrants to the agency that it will produce design documents that 

are complete and free from error (design-builder takes the risk). Standard 

specifications are used In CM/GC contract documents.
283

Table 14-1 Items to Include in a Driven Pile Specification

Category Item

Pile Materials Material type, grade, and strength. 

Coating details.

Transportation and handling.

Driving System and Equipment Hammer.

Hammer and pile cushions.

Helmet and inserts. 

Pile leads. 

Followers.

Predrilling equipment. 

Jetting equipment. 

Spudding equipment. 

Installation Issues Driving sequence.

Pile location tolerances.

Pile alignment tolerances.

Pile shoe or toe protection requirement.

Pile splices.

Pile cutoff.

Pile cap connection.

Pile heave.

Pile rejection criteria. 

Resistance Verification Static load testing.

Dynamic testing.

Rapid load testing.

Wave equation analysis.

Dynamic formulas.

Basis of Payment Method of measurement.

Payment items.
284

14.2 GENERIC DRIVEN PILE SPECIFICATION

SECTION CONTENTS PAGE

Section X.01. DESCRIPTION 286 

Section X.02. SUBMITTALS AND APPROVALS 287 

A. Pile Installation Plan 287 

B. Pile Driving Equipment Approval by Wave Equation 288 

C. Pile Driving Equipment Approval by Alternate Method 294 

Section X.03. MATERIALS 295 

Section X.04. DRIVING EQUIPMENT AND APPURTENANCES 296 

A. Pile Hammers 296 

B. Drive System Components and Accessories 298 

1. Hammer Cushion 298 

2. Helmet 298 

3. Pile Cushion 299

4. Leads 300 

5. Followers 300 

6. Jets 301 

7. Predrilling Equipment 302 

8. Spuds 302 

Section X.05. DETERMINATION OF NOMINAL RESISTANCE 303 

A. Axial Compression Resistance 303 

1. Static Load Tests 303 

2. Dynamic Testing with Signal Matching 305 

3. Rapid Load Tests 309

4. Wave Equation Analysis 311 

5. Dynamic Formula 311 

B. Axial Tension Resistance 313

1. Static Load Tests 313 

2. Dynamic Testing with Signal Matching 314 

C. Lateral Resistance 315 

1. Static Load Tests 315 

Section X.06. PREPARATION AND DRIVING 316 

A. Site Work 316 

1. Excavation and Fill Placement 316 

2. Predrilling for Driving 317 
285

3. Predrilling through Embankments 317 

B. Preparation of Piling 318

1. Pile Heads 318

2. Collars for Timber Piles 318

3. Pile Shoes and Closure Plates 318

4. Pile Marking for Prior to Driving 318

C. Driving 319

1. Test, Probe, and Indicator Piles 319

2. Production Piles 320 

3. Driving Stress 321 

4. Installation Sequence 321 

5. Location, Alignment and Cutoff Tolerance 321 

6. Heaved Piles 322 

7. Obstructions 322 

8. Practical and Absolute Refusal 322 

9. Pile Splices 323 

10. Pile Cutoffs 324

11. Unsatisfactory Piles 324 

Section X.07. METHOD OF MEASUREMENT 325 

A. Mobilization of Pile Driving Equipment 325 

B. Piles Furnished 326 

C. Piles Driven 326

D. Test Piles Furnished 326

E. Test Piles Driven 326

F. Static Pile Load Test 327 

G. Dynamic Pile Load Test 327 

H. Rapid Pile Load Test 327 

I. Lateral Load Test 328

J. Splices 328

K. Pile Shoes 328

L. Predrilling 328

M. Jetting 329

N. Pile Cutoff 329

O. Spudding 329

P. Delays, Downtime, Rig Moves 329

Section X.08. BASIS OF PAYMENT 329
286

14.2.1 SECTION X.01 DESCRIPTION

This item shall consist of furnishing and driving foundation piles of the type and 

dimensions designated, including cutting off or building up foundation piles when 

required. Piling shall conform to and be installed in accordance with these 

specifications, and at the location, and to the elevation, penetration depth and/or 

nominal resistance shown on the plans, or as directed by the Engineer.

Except when test piles are required, the Contractor shall furnish the piles in 

accordance with the dimensions shown in the contract documents. When test piles 

are required, the pile lengths shown on the plans are for estimating purposes only 

and the actual lengths to be furnished for production piles will be determined by the 

Engineer after the test piles have been driven and tested. The lengths given in the 

order list will be based on the lengths which are assumed after cutoff to remain in 

the completed structure. The Contractor shall, without added compensation, 

increase the lengths to provide for fresh heading and for such additional length as 

may be necessary to suit the Contractor's method of operation.

Where required by the contract documents, the Contractor shall perform the static 

and/or dynamic and/or rapid load tests of the type and quantity specified at the 

locations indicated on the plans. 

Commentary:

The objective of this specification is to provide criteria by which the Owner 

can assure that designated piles are properly installed and the Contractor can 

expect equitable compensation for work performed. The Owner's 

responsibility is to estimate the pile lengths required for the nominal 

resistances. Pile lengths should be estimated based on subsurface 

explorations, testing and analysis which are completed during the design 

phase, and expected variability in the subsurface conditions. Pile contractors 

who enter contractual agreements to install piles for an owner should not be 

held accountable or indirectly penalized for inaccuracies in estimated lengths. 

The Contractor's responsibility is to provide and install designated piles, 

undamaged, to the length specified by the Owner. This work is usually 

accomplished within an established framework of restrictions necessary to 

insure a "good" pile foundation. The price bid for this item of work will reflect 

the Contractor's estimate of both actual cost to perform the work and 

perceived risk.
287

14.2.2 SECTION X.02 SUBMITTALS AND APPROVALS

A) Pile Installation Plan:

A Pile Driving Installation Plan shall be prepared by the Contractor and submitted to 

the Engineer no later than 30 days before driving the first pile. The Pile Driving 

Installation Plan shall include the following:

1. List and size of proposed equipment including cranes, barges, pile driving 

equipment, jetting equipment, compressors, and predrilling equipment. 

Manufacturer’s data sheets on hammers should be included.

2. Methods to determine hammer energy in the field for determination of nominal 

resistance. Include in the submittal necessary charts and recent calibrations for 

any pressure measuring equipment.

3. Detailed drawings of any proposed followers.

4. Detailed drawings of any templates.

5. Details of proposed load test equipment, all load test instrumentation, and load 

test procedures. Submitted load test program information should also include the 

details and arranged of the reaction frame and reaction piles as well as recent 

calibrations of jacks, required load cells and all monitoring equipment.

6. Sequence of driving of piles for each different configuration of pile layout.

7. Proposed schedule for test pile program and production pile driving.

8. Details of proposed means and procedures to document the preconstruction 

condition of existing nearby structures and utilities as well as proposed 

procedures for their protection and monitoring during the contract period.

9. Required shop drawings for piles, cofferdams, etc.

10.Methods and equipment proposed to prevent movement of piles during 

placement and compaction of fill within 15 feet of the piles.
288

11.Proposed method for placing steel reinforcement and concrete in concrete filled 

pipe piles.

12.Methods to prevent deflection of battered piles due to their own weight and to 

maintain their as-driven position until casting of the pile cap is complete.

13.Proposed pile splice locations and details of any mechanical or proprietary 

splices to be used.

B. Pile Driving Equipment Approval by Wave Equation:

All pile driving equipment furnished by the Contractor shall be subject to the 

approval of the Engineer. All pile driving equipment should be sized such that the 

project piles can be driven with reasonable effort to the estimated contract lengths 

without damage. Approval of pile driving equipment by the Engineer will be based 

on wave equation analysis unless a dynamic formula has been specified for nominal 

resistance verification in the field. In no case shall the driving equipment be 

transported to the project site until approval of the Engineer is received in writing. 

The Contractor shall submit a completed Pile Driving and Equipment Data Form 

(shown in Figure 14-1) to the Engineer for approval as part of the pile installation 

plan at least 30 days prior to the start of pile driving. If a follower is to be used, 

detailed drawings of the follower shall be included as part of this submittal.

Commentary:

Use of wave equation analysis for approval of driving equipment can 

substantially reduce pile driving costs and pile driving claims by checking that 

the equipment mobilized to the project can drive the pile to the required 

penetration depth without damage. Agencies should encourage Contractors 

to use wave equation analysis to select the optimum hammer for each 

project. In cases where disputes arise over rejection of pile driving 

equipment, the Engineer should request the Contractor to submit proof of the 

adequacy of the pile driving equipment. Proof should consist of, but not be 

limited to, a wave equation analysis of the proposed driving equipment 

performed by a registered professional engineer. All costs of this submission

shall be the responsibility of the Contractor. The Pile and Driving Equipment

Data Form should be submitted for approval even if wave equation analysis 

will not be used for hammer approval. The approved form should be used by 
289

 

Figure 14-1 Drive system submittal form.
290

the pile inspector to check the proposed hammer and drive system 

components are as furnished and are maintained during the driving operation.

The criteria, which the Engineer will use to evaluate the pile driving equipment from 

the wave equation results, consists of both the required number of hammer blows 

per foot of penetration as well as the pile driving stresses at the required nominal 

resistance. The required penetration resistance (blow count) indicated by the wave 

equation at the required nominal resistance shall be between 30 and 96 blows per 

foot for the driving equipment to be acceptable.

Commentary:

Practical refusal is defined later in this generic specification as 10 blows per 

inch. Therefore, the upper limit of the penetration resistance for hammer 

approval should be less than the criteria for practical refusal. Otherwise, 

slight variations in assumed soil behavior or hammer performance will result 

in refusal driving conditions occurring prior to achieving the nominal 

resistance with an approved driving system.

In addition, the pile driving stresses indicated by the wave equation analysis of the 

proposed driving equipment shall not exceed material specific limits for the driving 

system to be acceptable. The AASHTO (2014) resistance factor for driven pile 

drivability analysis, φda, is 1.0 for all pile types with the exception of timber piles 

where it is 1.15. For steel piles, maximum compressive driving stresses shall not 

exceed the resistance factor, φda, times 90 percent of the minimum yield strength of 

the pile material. For prestressed concrete piles in normal environments, tensile 

stresses shall not exceed φda times 0.095 multiplied by the square root of the 

concrete compressive strength, f'c, plus the effective prestress value, fpe (with both f'c

and fpe in ksi). For prestressed concrete piles in severe corrosive environments, 

tensile stresses shall not exceed φda times fpe. Compressive stresses for 

prestressed concrete piles shall not exceed φda times 85 percent of the compressive 

strength minus the effective prestress value, (i.e. 0.85 f'c - fpe). For timber piles, the 

compressive driving stress shall not exceed φda times 2.6 times the reference design 

value of wood in compression parallel to the grain, Fco, as listed on the plans or 

provided in AASHTO (2014) Table 8.4.1.4-1. 

The Contractor will be notified of the acceptance or rejection of the driving system 

within 14 calendar days of the Engineer's receipt of the Pile and Driving Equipment 

Data Form. If the wave equation analyses show that either pile damage or inability 

to drive the pile with a reasonable driving resistance to the desired nominal 
291

resistance will result from the Contractor's proposed equipment or methods, the 

Contractor shall modify or replace the proposed methods or equipment at his 

expense until subsequent wave equation analyses indicate the piles can be 

reasonably driven to the desired nominal resistance, without damage. The Engineer 

will notify the Contractor of the acceptance or rejection of the revised driving system 

within 7 calendar days of receipt of a revised Pile and Driving Equipment Data Form.

During pile driving operations, the Contractor shall use the approved system. No 

variations in the driving system will be permitted without the Engineer's written 

approval. Any change in the driving system will only be considered after the 

Contractor has submitted the necessary information for a revised wave equation 

analysis. The Contractor will be notified of the acceptance or rejection of the driving 

system changes within 7 calendar days of the Engineer's receipt of the requested 

change. The time required for submission, review, and approval of a revised driving 

system shall not constitute the basis for a contract time extension to the Contractor.

Commentary:

The nominal driving resistance, Rndr, is the soil resistance which must be 

overcome (including resistance from unsuitable layers, liquefiable soils, and 

scour zone soils) to reach the pile penetration depth where the factored 

resistance can be achieved with an appropriate resistance factor based on 

the nominal resistance verification method used in the field. The resistance 

factor depends on the reliability of the resistance determination method as 

well as, for some methods, the number of tests performed. Table 14-2

provides a summary of the resistance determination methods and their 

associated resistance factors. AASHTO does not provide a recommended 

resistance factor for rapid load test methods as discussed in Section 11.4.7.

The nominal driving resistance is affected by:

1. The resistance in unsuitable soil support layers overlying suitable 

support layers. 

2. Minimum penetration requirements.

3. Temporary loss or increase in soil strength due to driving operations. 

4. Pile installation methods which alter the in place soil resistance such as 

jetting, predrilling, etc.
292

The designer must estimate the nominal driving resistance. Only on the most 

routine pile projects will the nominal driving resistance be equal to nominal 

resistance (i.e. the nominal resistance is the factored resistance divided by 

the resistance factor). More typically, piles are used to penetrate upper soil 

layers which are unsuitable for load support due to either poor soil 

characteristics, or future loss of load support by scour or liquefaction. In such 

cases, resistance in the unsuitable layers is not considered in determining the 

pile penetration necessary to support the factored resistance. However, the 

estimated nominal driving resistance must include the resistance encountered

in penetrating those unsuitable support layers, in addition to the nominal 

resistance.

The nominal driving resistance must be shown on the contract documents to 

permit the Contractor to properly size the driving equipment and the Engineer 

to judge the acceptability of the Contractor's driving equipment. Optimum pile 

installation generally occurs when the nominal driving resistance is obtained 

with a driving effort below the point of maximum curvature (typically around 

84 to 96 blows per foot) of the wave equation bearing graph. Larger 

penetration resistance result in negligible pile penetration per blow and 

generally inefficient driving conditions. Excessive driving resistances can also 

result in damage to the pile or the driving system.
293

Table 14-2 Resistance Factors for Field Determination Methods 

(after AASHTO 2014)

Nominal Resistance of Single Pile Using a Dynamic 

Analysis and Static Load Test Method

AASHTO 

Resistance 

Factor, φdyn

Compression Resistance: Driving criteria established by 

successful static load test of at least one pile per site 

condition and dynamic testing with signal matching of at 

least two piles per site condition, but no less than 2% of 

the production piles.

0.80

Compression Resistance: Driving criteria established by 

successful static load test of at least one pile per site 

condition without dynamic testing.

0.75

Compression Resistance: Driving criteria established by 

dynamic testing with signal matching conducted on 100% 

of production piles. 

0.75

Compression Resistance: Driving criteria established by 

dynamic testing with signal matching of at least two piles 

per site condition, but no less than 2% of the production 

piles.

0.65

Compression Resistance: Wave equation analysis, 

without pile dynamic measurements or load test, at end 

of drive conditions only.

0.50

Compression Resistance: FHWA Modified Gates 

dynamic pile formula (End of Drive condition only). 0.40

Compression Resistance: Engineering News dynamic 

pile formula (End of Drive condition only). 0.10

Tension Resistance: Static load test. 0.60

Tension Resistance: Dynamic testing with signal 

matching. 0.50
294

C) Pile Driving Equipment Approval by Alternate Method:

An alternate method of driving equipment approval will be used when the contract 

documents state that nominal resistance verification in the field will be determined 

from a dynamic formula. The alternate approval method requires that the energy of 

the driving equipment be rated by the manufacturer at or above the appropriate 

minimum energy level in Table 14-3 corresponding to the nominal resistance shown 

on the plans. The penetration resistance required by the dynamic formula for the 

submitted hammer and required nominal resistance should not exceed 10 blows per 

inch.

During pile driving operations, the Contractor shall use the approved system. If the 

Engineer determines the Contractor's hammer is unable to transfer sufficient energy 

to the pile, the hammer shall be removed from service until repaired to the 

satisfaction of the Engineer. No variations in the driving system will be permitted 

without the Engineer's written approval. Any changes in the driving system will be 

considered only after the Contractor has submitted a new Pile and Driving 

Equipment Data form. The Contractor will be notified of the acceptance or rejection 

of the proposed change in driving equipment within 7 calendar days of the 

Engineer's receipt of the form.

Table 14-3 Alternate Approval Method Minimum Pile Hammer Requirements

Nominal Resistance

(kips) 

Minimum Manufacturers Rated

Hammer Energy (ft-lbs) *

180 and under -----

181 to 300 -----

301 to 420 -----

421 to 540 -----

541 to 600 -----

Over 600 -----

 * See commentary.

Commentary:

A table of the minimum rated hammer energy vs. nominal resistance should 

be developed using wave equation analyses of commonly available driving 

systems for the pile types, pile lengths, and pile loads routinely used by the 

specific agency. These analyses should model the typical soil and pile 
295

installation conditions. The wave equation results should be evaluated for 

driving stress levels and penetration resistance (blow count) to determine 

which hammer energies are too large (driving stress problems or penetration 

resistance at nominal resistance less than 30 blows/ft) and which energies 

are too small (penetration resistance at nominal resistance greater than 

practical refusal value of 120 blows /ft).

Once the specific table of energy values has been developed, it should only 

be considered for routine projects in uniform soil conditions. Projects 

involving long piles or large nominal resistances relative to the factored load 

(such as piles subject to significant scour depths or piles to be driven through 

embankments) should use project specific wave equation analyses to 

establish minimum driving equipment requirements. Piles to soft and hard 

rock should also be evaluated by wave equation analysis to reduce the risk of 

pile damage from too large a hammer.

14.2.3 SECTION X.03 MATERIALS

Materials shall meet the requirements in the following Subsections of Section X.03

Materials:

Portland Cement Concrete

Reinforcing Steel

Prestressing Strands / Post-Tensioning Tendons

Structural Steel

Castings for Pile Shoes

Steel Shells for Cast in Place Piles

Timber Piles

Timber Preservative and Treatment 

Protective Coatings

Commentary:

The appropriate sections of each agency's standard specifications should be 

included under the X.03 Materials. A generic materials section cannot be 

provided herein, considering the vast combinations of materials used in piling 

operations and the varying control methods used by individual agencies. The 

above list contains the common material components. Additions or deletions 

may be required to this list based on the content of individual agency 
296

standard specifications and the pile type specified. The 3rd Edition of the 

AASHTO LRFD Bridge Construction Specifications contains additional text 

and commentary on pile material topics which should be reviewed and used 

to update agency standard specifications as appropriate.

14.2.4 SECTION X.04 DRIVING EQUIPMENT AND APPURTENANCES

A) Pile Hammers:

Piles may be driven with air, steam, diesel, or hydraulic hammers. Drop hammers, if 

specifically permitted in the contract, shall not be used for concrete piles or for piles 

whose required nominal resistance exceeds 120 kips. When drop hammers are 

permitted, the ram shall have a weight not less than 2.0 kips and the height of the 

drop shall not exceed 12.0 feet. In no case shall the ram weight of the drop hammer 

be less than the combined weight of helmet and pile. All drop hammers shall be 

equipped with hammer guides to insure concentric impact on the helmet.

Air/steam hammers shall be operated and maintained within the manufacturer’s 

specified ranges. The plant and equipment furnished for air/steam hammers shall 

have sufficient capacity to maintain at the hammer, under working conditions, the 

volume and pressure specified by the manufacturer. The hose connecting the 

compressor or boiler with the hammer shall be at least the minimum size 

recommended by the hammer manufacturer. The plant and equipment shall be 

equipped with accurate pressure gauges which are easily accessible to the 

Engineer. The weight of the striking parts of air/steam hammers shall not be less 

than one third the weight of helmet and pile being driven, and in no case shall the 

striking parts weigh less than 2.75 kips. If a wave equation analysis is used for 

hammer approval the minimum ram weight requirements shall not apply.

Open end (single acting) diesel hammers shall be equipped with a device such as 

rings on the ram to permit the Engineer to visually determine hammer stroke at all 

times during pile driving operations. Also, the Contractor shall provide the Engineer 

a chart from the hammer manufacturer equating stroke and blows per minute for the 

open end diesel hammer to be used. For open end diesel hammers, the contractor 

shall provide and maintain in working order for the Engineer’s use, an approved 

device to automatically determine and display ram stroke.

Closed end (double acting) diesel hammers shall be equipped with a bounce 

chamber pressure gauge, in good working order, mounted near ground level so as 
297

to be easily read by the Engineer. Also, the Contractor shall provide the Engineer a 

chart, calibrated to actual hammer performance within 60 days of use, equating 

bounce chamber pressure to either equivalent energy or stroke for the closed end 

diesel hammer to be used.

Hydraulic hammers shall be equipped with a system for measuring and immediately 

displaying in the field, the kinetic energy or ram impact velocity. The system shall be 

maintained in good working order and operational at all times piles are driven.

Vibratory hammers, when permitted for installing production piles, shall be used only 

after the pile toe elevation for the nominal resistance is established by load testing 

and/or from test piles restruck with an impact hammer. The Contractor shall 

perform, at his cost, the load tests and/or extra work required by the Engineer as 

needed for approval of the vibratory hammer use. Installation of production piles 

with vibratory hammers shall be controlled according to power consumption, rate of 

penetration, specified toe elevation, or other means acceptable to the Engineer 

which assures the nominal resistance equals or exceeds the nominal resistance of 

the test pile. In addition, the first of every 10 piles installed with a vibratory hammer 

shall be restruck with an impact hammer of suitable energy to verify the nominal 

resistance before driving the remaining piles.

Commentary:

Pile inspectors frequently do not possess adequate knowledge or technical

information concerning even the most basic details of the Contractor's 

hammer. Chapters 15 and 18 provide information on driving equipment and

monitoring. Agencies and contractors should also provide pile “inspectors” 

with basic manuals such as FHWA/RD 86/160 "The Performance of Pile 

Driving Systems: "Inspections Manual" or "Inspectors Manual for Pile 

Foundations" and "A Pile Inspectors Guide to Hammers, Second Edition" 

available from the Deep Foundation Institute, 120 Charlotte Place, Englewood 

Cliffs, NJ 07632.

On large projects or on projects requiring large pile resistances, specifications 

should consider requiring kinetic energy readout devices for hammers as 

described in Section 15.19 of Chapter 15. Several manufacturers can equip 

their hammers with these devices when requested. Any existing hammer can 

also be retrofitted with a kinetic energy readout device. These devices allow 

improved quality assurance and can detect changes in hammer performance 

over time that may necessitate adjustment to the pile installation criterion.
298

At present no formula exists to reliably predict the nominal resistance of piles 

driven with vibratory hammers. Until reliable procedures are developed for 

vibratory installation, special precautions must be taken to insure foundation 

piles installed with vibratory hammers have both adequate nominal resistance 

and structural integrity. As discussed in Section 7.10.5, the use of vibratory 

hammers may also affect the shaft resistance that develops on the pile. On 

critical projects, “owners” should consider the use of dynamic testing during 

restrike to substantiate pile nominal resistance and integrity.

B) Drive System Components and Accessories:

1. Hammer Cushion: Impact pile driving equipment designed to be used with a 

hammer cushion shall be equipped with a suitable thickness of hammer cushion 

material to prevent damage to the hammer or pile and to insure uniform driving 

behavior. Hammer cushions shall be made of durable manufactured materials, 

provided in accordance with the hammer manufacturer's guidelines. Wood, wire 

rope, and asbestos hammer cushions are specifically disallowed and shall not be 

used. A striker plate, as recommended by the hammer manufacturer, shall be 

placed directly above the hammer cushion to insure uniform compression of the 

cushion material. The hammer cushion shall be removed from the helmet and 

inspected in the presence of the Engineer when beginning pile driving at each 

structure or after each 100 hours of pile driving, whichever is less. Any reduction of 

hammer cushion thickness exceeding 25 percent of the original thickness shall be 

replaced by the Contractor before driving is permitted to continue.

Commentary:

For hammers requiring cushion material, mandatory use of a durable hammer 

cushion material that will retain uniform properties during driving is necessary to 

accurately relate penetration resistance (blow count) to nominal pile resistance. 

Non-durable materials which deteriorate during driving cause erratic estimates 

of nominal resistance and, if allowed to dissolve, result in damage to the pile or 

driving system. 

2. Helmet: Piles driven with impact hammers require an adequate helmet or drive 

head to distribute the hammer blow to the pile head. The surface of the helmet in 

contact with the pile shall be plane and smooth and shall be aligned parallel with the 

hammer base and the pile head. The helmet shall be guided by the leads and not 

be free swinging. The helmet shall fit around the pile head in such a manner as to 

prevent transfer of torsional forces during driving, while maintaining proper alignment 
299

of hammer and pile. An insert may be used with a helmet to adapt the helmet to 

different types or sizes of piles.

For steel and timber piling, the pile heads shall be cut squarely and a helmet, as 

recommended by the hammer manufacturer, shall be provided to hold the axis of the 

pile in line with the axis of the hammer.

For timber piles, the least inside helmet or hammer base horizontal dimension shall 

not exceed the pile head diameter by more than 2.0 inches. If the timber pile 

diameter slightly exceeds the least helmet or hammer base dimension, the pile head 

shall be trimmed to fit the helmet.

For precast concrete and prestressed concrete piles, the pile head shall be plane 

and perpendicular to the longitudinal axis of the pile to prevent eccentric impacts 

from the helmet.

For special types of piles, appropriate helmets, mandrels or other devices shall be 

provided in accordance with the manufacturer's recommendations so that the piles 

may be driven without damage.

3. Pile Cushion: The heads of concrete piles shall each be protected by a pile 

cushion having the same cross sectional area as the pile top. Pile cushions shall be 

made of plywood, hardwood, or composite plywood and hardwood materials.

The minimum pile cushion thickness placed on the pile head prior to driving shall be 

determined by wave equation analysis so that driving stress limits are not exceeded. 

If a dynamic formula is used, the minimum pile cushion thickness shall be at least 4 

inches.

A new pile cushion shall be provided for each pile. In addition, the pile cushion shall 

be replaced during the driving of any pile if the cushion is compressed more than 

one-half the original thickness or it begins to burn. Pile cushions shall be protected 

from the weather, and kept dry prior to use. Pile cushion shall not be soaked in any 

liquid unless approved by the Engineer. The use of manufactured pile cushion 

materials in lieu of a wood pile cushion shall be evaluated on a case by case basis.

A used pile cushion in good condition shall be used for restrike tests. The used pile 

cushion shall be the same cushion from the end of initial driving unless that pile 

cushion condition has deteriorated. If the original pile cushion has deteriorated, 
300

another used pile cushion of similar thickness as the original cushion at the end of 

drive shall be used.

Commentary:

A pile cushion is only needed for the protection of concrete piles. If the wave 

equation analysis of the Contractor's hammer indicates tension stresses 

exceed specification limits, the pile cushion may need to be substantially 

thicker than 4 inches. Pile cushion thicknesses greater than 18 inches have 

been used to mitigate tension stresses. Compressive stresses at the pile 

head can generally be controlled with a relatively thin pile cushion. However, 

wood pile cushions may become overly compressed and hard after about 

1000 to 1500 hammer blows. Conversely, cushions exposed to less than 50 

blows are generally not suitable for restrikes.

4. Leads: Piles shall be supported in line and position with leads while being 

driven. Pile driver leads shall be constructed in a manner that affords freedom of 

movement of the hammer while maintaining alignment of the hammer and the pile to 

insure concentric impact for each blow.

Leads may be either fixed or swinging type. Swinging leads shall be adequately 

embedded in the ground or the pile constrained in a structural frame such as a 

template to maintain alignment and location tolerances. Swinging leads shall be 

fitted with a pile gate at the bottom of the leads unless used with a template. Leads 

shall be of sufficient length to make the use of a follower unnecessary. Leads used 

for driving batter piles shall be designed to permit and maintain proper alignment of 

the batter pile. A horizontal brace may be required between the crane and base of 

leads to maintain alignment and location tolerances in some batter pile installation 

conditions.

5. Followers: Followers shall only be used when approved in writing by the 

Engineer, or when specifically stated in the contract documents. When a follower is 

proposed, a wave equation analysis shall be used to evaluate the suitability of the 

proposed driving system. As a general guide, the cross sectional area of a steel 

follower when driving concrete piles should be at least 20 percent of the cross 

sectional area of the concrete pile. When driving steel or timber piles, the cross 

sectional area of the steel follower should have an impedance between 50 percent 

and 200 percent of the pile impedance.
301

The follower and pile shall be held and maintained in equal and proper alignment 

during driving. The follower shall be of such material and dimensions to permit the 

piles to be driven to the penetration depth determined necessary from the driving of 

the full length piles. The follower shall be designed with guides adapted to the leads 

that maintain the hammer, follower and pile in alignment during driving. The lower 

end of the follower shall be equipped with a helmet or follower-pile connection 

suitable for the pile type being driven.

The final position and alignment of the first two piles installed with followers in each 

substructure unit shall be verified to be in accordance with the location and 

alignment tolerances in Section X.06 C) 4 before additional piles are installed.

Commentary:

The use of a follower often causes substantial and erratic reductions in the 

hammer energy transmitted to the pile due to the follower flexibility, poor 

connection to the pile head, frequent misalignment, etc. Reliable correlations 

of penetration resistance with nominal resistance are very difficult when 

followers are used. Therefore, the nominal resistance of select follower 

driven piles should be checked with either a static load test or dynamic testing 

with signal matching. Severe problems with pile alignment and location 

frequently occur when driving batter piles with a follower in a cofferdam 

unless a multi-tier template is used.

6. Jets: Jetting shall only be permitted if approved in writing by the Engineer or 

when specifically stated in the contract documents. The contractor shall determine 

the number of jets and the volume and pressure of water at the jet nozzles to freely 

erode the material adjacent to the pile without affecting the lateral stability of the final 

in place pile.

The Contractor shall control and dispose of all jet water in a manner satisfactory to 

the engineer or as specified in the contract documents. If jetting is specified or 

approved by the engineer and the jetting is performed as specified or approved, the 

contractor shall not be held responsible for any damage to the site caused by the 

jetting operations. If jetting is performed for the contractor’s convenience, the 

contractor shall be responsible for all damage to the site caused by jetting 

operations. 

Jet pipes shall be removed when the bottom of the jet pipe is 5 feet above the 

minimum or prescribed toe elevation unless otherwise indicated by the contract 
302

documents or the Engineer. Following jet removal, the jetted pile shall be driven to 

the required nominal resistance with an impact hammer. If the required nominal 

resistance is not achieved at the prescribed toe elevation, the pile may be allowed to 

setup and the required nominal resistance determined through restriking. The jetting 

procedures should be reviewed by the Engineer and adjustments made if applicable 

so that the required nominal resistance can be achieved without restrike 

verifications.

When jetting is used, the Contractor shall submit details of the proposed jetting and 

pile driving plan. Where practical, all piles in a pile group shall be jetted to the 

required penetration depth before beginning pile driving. When large pile groups or 

pile spacing and batter make this impractical, restrike tests on a select number of 

previously driven piles shall be performed to check nominal resistance after jetting 

operations are completed.

7. Predrilling Equipment: When stated in the contract documents, the 

Contractor shall provide predrilling equipment to drill holes at pile locations of the 

size specified and to the depths shown in the contract documents or as approved in 

writing by the Engineer. If subsurface obstructions, such as boulders or rock layers 

are encountered, the diameter of the predrilled hole may be increased with 

Engineers approval to the least dimension adequate for pile installation.

Commentary: 

The appropriate diameter of the predrilled hole depends on the purpose of the 

predrilled hole. If predrilling is performed to minimize problems with 

maintaining alignment tolerances, or to mitigate heave or vibrations, predrilled 

holes are typically smaller than the diameter or diagonal of the pile. When 

predrilling is performed to penetrate through an embankment or to bypass 

obstructions, a larger predrilled hole with a diameter not more than the largest 

dimension of the pile plus 6 inches may be acceptable. In either case, the 

excavated zone surrounding the pile is generally backfilled with an approved 

material of sand, pea gravel, or grout after the pile is driven depending on 

design requirements.

8. Spuds: When stated in the contract documents or approved by the Engineer, 

the Contractor shall provide spudding equipment to displace obstructions, debris, or 

unsuitable materials at pile locations. The spudding equipment shall create an 

opening through the material of the size specified and to the depths shown in the 

contract documents or as approved by the Engineer in writing.
303

14.2.5 SECTION X.05 DETERMINATION OF NOMINAL RESISTANCE

A) Axial Compression Resistance:

The nominal resistance of piles in axial compression shall be determined by the 

Engineer based on one of the methods listed below.

1. Static Load Tests: Compression load tests shall be performed by 

procedures set forth in ASTM D1143 using the quick load test method, except that 

the test shall be taken to geotechnical plunging failure or the capacity of the loading 

system. Testing equipment and measuring systems shall conform to ASTM D1143, 

except that the loading system shall be capable of applying 150 percent of the 

nominal resistance. A load cell and spherical bearing plate shall be used.

The Contractor shall submit to the Engineer for approval detailed plans prepared by 

a licensed professional engineer of the proposed loading apparatus. The submittal 

shall include calibrations for the hydraulic jack, load cell, and pressure gage 

conducted within 30 days of the load test. If requested by the Engineer, the jack, 

load cell, and pressure gage shall be recalibrated after the load test.

The loading apparatus shall be constructed to allow the various increments of the 

load to be placed gradually, without causing vibration to the test pile. When the 

approved method requires the use of reaction piles, the reaction piles shall be of the 

same type and diameter as the production piles. Reaction piles shall be surveyed 

and monitored for upward movement during the load test. Reaction piles driven at 

production pile locations that have a permanent upward movement of 0.25 inches or 

more upon completion of the load test shall be redriven. Timber or tapered piles 

installed in permanent locations shall not be used as reaction piles.

While performing the load test, the contractor shall provide safety equipment and 

employ adequate safety procedures. Adequate support for the load test plates, jack, 

and ancillary devices shall be provided to prevent them from falling in the event of a 

release of load due to hydraulic failure, test pile failure, or other cause.

The nominal geotechnical resistance or failure load of a pile statically tested in axial 

compression is defined by the pile head movement under load. For piles 24 inches 

or less in diameter or width, the failure load is the pile head load which produces a 

measured movement of the pile head equal to:





 



 ++∆=

120

15.0 bsf Eq. 14-1 
304

Where:

sf = pile head movement at failure (inches).

Δ = elastic deformation of total pile length (inches).

b = pile diameter or width of side for square piles (inches).

For piles larger than 36 inches in diameter, additional pile toe movement is 

necessary to develop the toe resistance. For these larger diameter piles, the 

nominal geotechnical resistance or failure load can be defined as the load which 

produces movement at the pile head equal to:

bs 30/ f +∆= Eq. 14-2 

Where: 

sf = pile head movement at failure (inches).

Δ = elastic deformation of total pile length (inches).

 b = pile diameter or width of side for square piles (inches).

For piles greater than 24 inches in diameter but less than 36 inches in diameter, 

linear interpolation should be performed between Eq. 14-1 and 14-2.

The top elevation of the test pile shall be determined immediately after driving and 

again just before load testing to check for heave. If more than 1⁄4 inch of heave 

occurs the load test pile may require re-driving before the load test is performed.

Unless otherwise specified in the contract documents or by the Engineer, the static 

load test shall not be performed sooner than 5 days after the test pile or any reaction 

piles were driven.

On completion of the load testing, any test or reaction piling not a part of the finished 

structure shall be removed or cut off at least 1 foot below either the bottom of footing 

or the finished ground elevation, if not located within the footing area.

Commentary: 

The nominal resistance may increase (soil setup) or decrease (relaxation) 

after the end of driving. Therefore, it is essential that static load testing be 

performed after equilibrium conditions in the soil have re-established. Static 

load tests performed before equilibrium conditions have re-established will 

underestimate the long term nominal resistance in soil setup conditions and 

overestimate the long term nominal resistance in relaxation cases. For piles 
305

in clays, specifications should require at least 2 weeks or longer to elapse 

between driving and load testing. In sandy silts and sands, 5 days to a week 

is usually sufficient. Load testing of piles driven into shales should also be 

delayed for at least 2 weeks after driving. Additional discussion on time 

dependent changes in nominal resistance may be found in Section 7.2.4. 

Each static load test pile should determine the load transferred to the pile toe. 

Instrumentation commonly consists of strain gages and/or telltale rods 

mounted at varying depths above the pile toe. Also, a load cell and spherical 

bearing plate should be mounted between the load frame and the pile head to 

verify the readings from the hydraulic jack pressure gauge. Due to jack ram 

friction, loads indicated by a jack pressure gauge are commonly 10 percent to 

20 percent higher than the actual load imposed on the pile.

If the static load tests are to be performed by an independent firm retained by 

the Contractor and not by the Engineer, an additional specification section 

detailing the complete load test instrumentation monitoring requirements as 

well as the report submission requirements for the load test results and result 

interpretation must be added. A corresponding pay item must then be added 

to this specification for load test reporting. Alternatively, the report 

requirements can be described herein and then included as part of the static 

load test pay item.

When static load tests are used to control production pile driving, the time 

required to analyze and/or review the load test results as well as to establish 

driving criteria should be specified so that the delay time is clearly identified. 

Static load testing is discussed in greater detail in Chapter 9 of this manual.

2. Dynamic Testing with Signal Matching: Dynamic measurements shall be 

obtained using dynamic test processing equipment, calibrated transducers, and 

procedures set forth in ASTM D4945. The measurements will be taken by a 

qualified engineer during the driving of piles designated as dynamic test piles. 

Signal matching analysis shall be performed on representative data collected at the 

end of initial driving and at the beginning of each restrike events. Additional signal 

matching analysis may be performed as determined by the Engineer.

Commentary:

This section on dynamic testing covers only the Contractor's activities as they 

relate to the dynamic tests. If the dynamic tests are to be performed by an 
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independent firm retained by the Contractor and not by the Engineer, an 

additional specification section detailing the dynamic test analysis and 

reporting requirements must be added. Merely referencing the ASTM D4945 

standard is insufficient. ASTM D4945 does not specify signal matching 

requirements or their frequency; it does not specify if, how, and by whom, 

driving criteria are established; nor does it identify what substructure locations

are covered by the criteria.

Dynamic testing personnel should have attained an appropriate level of 

expertise (Expert, Master, Advanced, Intermediate, Basic, or Provisional) on 

the “Dynamic Measurement and Analysis Proficiency Test” sponsored by the 

Pile Driving Contractors Association (PDCA) and Pile Dynamics, Inc. (PDI) for 

providers of dynamic testing services. Dynamic testing methods are 

discussed in Chapter 10.

Whenever static load tests are specified, dynamic tests are recommended to 

be performed on at least half the reaction piles prior to driving the static load 

test pile as well as on the static load test pile when it is driven. The dynamic 

test results are used both to confirm that the desired nominal resistance can 

be attained at the estimated static load test pile penetration depth and to fine 

tune the dynamic test procedures for site soil conditions. Dynamic monitoring 

of the static load test pile during restrike after completion of the static load test 

is also highly recommended. This restrike test allows correlation of static test 

results with dynamic test results. Signal matching analysis of dynamic test 

data is required for nominal resistance determination per AASHTO (2014) 

and also assists in quantifying the dynamic soil parameters, soil quake and 

damping, for the site.

When dynamic tests are specified on production piles, the first pile driven in 

each substructure foundation is typically tested. The total number of dynamic 

tests performed will vary from two piles per site condition, but not less than 

2% of production piles, to 100% of the production piles. The number of 

dynamic tests required per AASHTO depends on the variability of the site 

conditions as well as the resistance factor selected for design verification. 

Additional discussion on the number of piles dynamically tested and the 

associated resistance factor can be found in Section 10.3.

Prior to placement in the leads, the Contractor shall make diametrically opposite 

faces of each pile to be dynamically tested available for predrilling the 

instrumentation attachment holes. The dynamic testing engineer will furnish the 
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equipment, materials, and labor necessary for drilling holes in the piles for mounting 

the instrumentation. The instruments will typically be attached 2 to 3 pile diameters 

below the head of the pile with bolts placed in masonry anchors for the concrete 

piles, or through drilled holes on the steel piles, or with wood screws for timber piles.

The Contractor shall provide the dynamic testing engineer reasonable means of 

access to the pile for attaching instruments after the pile is placed in the leads. A 

manlift or platform with minimum size of 4 feet x 4 feet designed to be raised to the 

top of the pile while the pile is located in the leads shall be provided and operated by 

the Contractor. Alternatively, Contractor’s personnel following the dynamic testing 

engineer’s instructions can attach the instruments to the pile after it is placed in the 

leads. For some pile types and project conditions, the dynamic testing engineer may 

also recommend instrument attachment to the pile prior to lifting. It is estimated that 

approximately 1 hour per pile will be needed for instrument attachment and removal.

If requested, the Contractor shall furnish electric power for the dynamic test 

equipment. The power supply at the outlet shall be 10 amp, 115 volt, 55-60 cycle, 

A.C. only. Field generators used as the power source shall be equipped with 

functioning meters for monitoring voltage and frequency levels.

For dynamic testing conducted from a barge or other difficult to access sites, the 

Contractor shall furnish a shelter to protect the dynamic test equipment from the 

elements. The shelter shall have a minimum floor size of 8 feet x 8 feet and 

minimum roof height of 6.5 feet. The inside temperature of the shelter shall be 

between 45 and 95 degrees and be located within 50 feet of the pile location.

With the dynamic testing equipment attached, the Contractor shall drive the pile to 

the design penetration depth or to a depth determined by the Engineer. The 

Engineer will use the nominal resistance estimates at the time of driving and/or 

restriking from dynamic test methods to determine the required pile penetration 

depth for the nominal resistance. The stresses in the piles will be monitored during 

driving so that the measured stresses do not exceed the specified material values in 

Section X.02 B). If necessary, the Contractor shall reduce the driving energy 

transmitted to the pile by using additional cushions or reducing the energy output of 

the hammer in order to maintain stresses below the values in Section X.02 B). If 

non axial driving is indicated by dynamic test equipment measurements, the 

Contractor shall immediately realign the driving system. 

The Contractor shall wait up to 24 hours (or a longer duration per Table 14-4 if 

specified in the contract documents) and restrike the dynamic load test pile with the 
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dynamic testing instruments attached. It is estimated that the Engineer will require 

approximately 30 minutes to reattach the instruments. A cold hammer shall not be 

used for the restrike. The hammer shall be warmed up before restrike begins by 

applying at least 20 blows to another pile or to timber mats placed on the ground.

The maximum amount of penetration required during restrike shall be 3 inches, or 

the maximum total number of hammer blows required will be 20, whichever occurs 

first. After restriking, the Engineer will either provide the cutoff elevation or specify 

additional pile penetration and testing.

Commentary:

A dynamic test often includes monitoring during both initial driving and during 

one restrike event within a specified time period. Alternatively, the initial 

driving and restrike test events can be individual items. Any long term restrike

tests after the initial restrike should be paid as a separate item unless the 

restrike schedule is specifically stated in the dynamic test specification.

The restrike time and frequency should be clearly stated in the specifications 

and should be based on the time dependent strength change characteristics 

of the soil. Table 14-4 provides restrike durations commonly used for various 

soil types.

Table 14-4 Common Time Delay for Restrike Based Upon Soil Type

Soil Type Time Delay Until Restrike

Clean Sands 1 Day

Silty Sands 2 Days

Sandy Silts 3-5 Days

Silty Clays 7-14 Days* 

Shales 10-14 Days* 

* Longer times sometimes required.

The restrike time interval is particularly important when dynamic testing is 

used for construction control. Specifying too short of a restrike time for 

friction piles in fine grained deposits may result in pile length overruns. 

However, it is sometimes difficult for long term restrikes to be accommodated 

in the construction schedule. In these cases, multiple restrikes are often 

specified on selected piles with shorter term restrikes at other locations.
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The time necessary to analyze the dynamic test results and provide driving 

criteria to the contractor once restrikes are completed should also be stated in 

the specifications. This is important when the testing is done by agency 

personnel or their consultants as well as when the testing firm is retained by 

the contractor. In cases where the testing is retained by the contractor, the 

time required for agency review of the test results and to provide driving 

criteria should be specified relative to the agency’s receiving the test results.

3. Rapid Load Tests: The nominal resistance shall be determined by a rapid 

load test performed in accordance with the procedures set forth in ASTM D7383. 

The rapid load test may be performed using either a combustion gas and reaction 

mass system or with a cushioned drop weight system. The peak force shall exceed 

the targeted nominal geotechnical resistance plus the dynamic soil resistance. The 

applied force pulse shall exceed 50% of the actual peak force for a time duration of 

4L/C and the static pre-load force for a time duration of at least 12L/C where L is the 

pile length in feet and C is the pile wave speed in ft/s. A time duration of less than 

12L/C is acceptable if additional force and movement measurements devices are 

used along the pile length in accordance with ASTM D7383.

The Contractor shall trim the head of the test pile flat, level, and perpendicular to the 

pile axis at the elevation directed by the rapid load testing agency. The area 

surrounding the test pile shall also be prepared in accordance with the requirements 

necessary to support the rapid load test device and conduct the test. Depending on 

the site conditions, this may necessitate site grading, use of crane pads, 

constructing a support frame, and/or other measures for positioning and supporting 

the rapid load test device.

The force shall be measured using a calibrated force transducer or load cell placed 

between the loading apparatus and the pile head. The load capacity of the 

transducer shall be at least 10% greater than the targeted peak force. The pile head 

movement shall be measured by one or more calibrated displacement transducers.

The nominal geotechnical resistance determined by a rapid load test shall be 

assessed by an approved interpretation procedure. The procedure shall identify the 

loading rate reduction factor used with the analysis method.

Commentary:

Section 11.4.1 through 11.4.5 describes nominal resistance interpretation 

methods for rapid load tests. The pile length and soil conditions are primary 
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factors in selecting the interpretation method. If additional force and 

movement devices are necessary for the interpretation method, time must be 

allotted for obtaining and installing this instrumentation in or on the test pile. 

A significant permanent pile head displacement is also needed for nominal 

resistance determination. The loading rate in a rapid load test also affects the 

nominal resistance and must be considered.

AASHTO (2014) does not include a resistance factor for rapid load tests. 

Hence, this must be stipulated by the designer whenever rapid load tests are 

specified. A discussion of resistance factors currently in use with rapid load 

tests is provided in Section 11.4.7. All rapid load tests are performed by 

independent testing firms. Therefore, the time required for agency review of 

the rapid load test results should be specified relative to when the agency’s 

receives the rapid load test results. A detailed discussion of rapid load test 

methods is provided in Chapter 11.

Unless otherwise specified in the contract documents or by the Engineer, the rapid 

load test shall not be performed sooner than 5 days after driving the test pile.

Commentary:

This specification addresses only the Contractor's activities as they relate to 

performing a rapid load test. If the rapid load tests are to be performed by an 

independent testing firm retained by the Contractor and not retained by the 

Engineer, an additional specification section detailing the complete rapid load 

test instrumentation monitoring requirements as well as the report submission 

requirements for the rapid load test results and result interpretation must be 

added. A corresponding pay item must then be added to this specification for 

rapid load test reporting. Alternatively, the report requirements can be 

described herein and then included as part of the rapid load test pay item.

The Contractor’s independent testing firm shall submit to the Engineer, for 

approval, detailed plans of the rapid load test equipment arrangement, 

proposed test procedure, and nominal resistance interpretation method. The 

submittal shall also include calibrations for the transducer or load cells and all 

additional instrumentation. All calibrations shall be within the calibration 

period specified within ASTM D7383.
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4. Wave Equation: The nominal resistance shall be determined by the 

Engineer based on a wave equation analysis. Piles shall be driven with the 

approved driving equipment to the ordered length or other lengths necessary to 

obtain the required nominal resistance. Jetting, predrilling, or other methods to 

facilitate pile penetration shall be modeled in the analysis if proposed and allowed by 

the contract documents. Adequate pile penetration depth for the nominal resistance 

shall be considered obtained when the wave equation penetration resistance is 

achieved within 5 feet of the estimated pile toe elevation. Piles not achieving the 

penetration resistance within this limit shall be driven to penetration depths 

established by the Engineer.

5. Dynamic Formula: The nominal resistance shall be determined by dynamic 

formula if the contract documents contain a provision that dynamic formula be used 

to establish driving criteria. Dynamic formulas should not be used if the required 

nominal resistance is greater than 600 kips. Formula results should not be 

considered applicable when the pile head is crushed, broomed, or damaged, or 

when a follower is used.

If a dynamic formula is used to establish driving criteria, piles shall be driven to a 

penetration depth necessary to obtain the nominal resistance according to the 

Modified Gates formula with specified units as follows:

100)10(log75.1 Rndr = Ed 10 Nb − Eq. 14-3 

Where: 

Rndr = nominal driving resistance (kips). 

Ed = developed hammer energy, (W)(h), during the observed set (ft-lbs).

W = ram weight (lbs).

h = average hammer stroke during set observation (ft).

Nb = number of hammer blows per inch (blows/in).

The number of hammer blows per foot of pile penetration required to obtain the 

nominal resistance shall be calculated as follows:

)10(12

x N ft = Eq. 14-4 

In which:

[( ) += 100 /( ]−1)75.1 Rx ndr Ed Eq. 14-5 
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Where: 

Nft = number of hammer blows for final foot of driving (blows/ft).

Rndr = nominal driving resistance (kips).

Ed = developed hammer energy, (W)(h), during the observed set (ft-lbs).

W = ram weight (lbs).

h = average hammer stroke during set observation (ft).

Commentary:

Additional dynamic formulas besides the FHWA Modified Gates formula were 

presented in Chapter 13. Other dynamic formulas may be used based on 

agency practice and local calibrations.

Driven piles should be monitored in terms of their nominal driving resistance. 

The nominal driving resistance at a given pile penetration depth reflects the 

total resistance mobilized by the pile. This may include resistance in soil 

deposits unsuited for long term load support, as well as suitable layers. 

Therefore, the penetration resistance (blow count) should be established for 

the nominal driving resistance that must be overcome in order to reach 

anticipated pile penetration depth. These nominal driving resistances are 

determined by static analysis procedures.

In the case of piles to be driven to a specified minimum pile toe elevation, the 

nominal driving resistance must be computed by static analysis to include the 

resistance of all soil layers penetrated by the pile above the minimum pile toe 

elevation as well as the toe resistance at that depth. The minimum pile toe 

elevation may have been specified for reasons other than axial compression 

resistance in order to meet lateral, uplift, or serviceability requirements. The 

nominal driving resistance is directly related to the maximum pile driving 

stress during installation. The driving stress is more critical than the stress 

imparted after installation by the factored design load.

Good piling practices dictate use of the wave equation in place of dynamic 

formulas. AASHTO design specifications require a wave equation drivability 

analysis be performed during the foundation design stage. The soil profile 

used in this design stage wave equation analysis can be easily re-used along 

with details on the Contractors proposed driving system to obtain a 

construction stage wave equation analysis that includes the penetration 

resistance and maximum pile stresses for the required nominal driving 
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resistance. FHWA recommends that all agencies use wave equation analysis 

with a goal of minimizing use of dynamic formulas on all pile projects. Wave 

equation analysis is discussed in greater detail in Chapter 12 of this manual.

B) Axial Tension Resistance:

The nominal resistance of piles in axial tension shall be determined by the Engineer 

based on one of the methods listed below.

1. Static Load Tests: Tension load tests shall be performed by procedures set 

forth in ASTM D3689 using the quick load test method, except that the test shall be 

taken to plunging failure or the capacity of the loading system. Testing equipment 

and measuring systems shall conform to ASTM D3689, except that the loading 

system shall be capable of applying 150 percent of the anticipated nominal 

resistance.

The Contractor shall submit to the Engineer for approval detailed plans prepared by 

a licensed professional engineer of the proposed loading apparatus. The submittal 

shall also include calibrations for the hydraulic jack, load cell, and pressure gage 

conducted within 30 days of the load test. If requested by the Engineer, the jack, 

load cell, and pressure gage shall be recalibrated after the load test.

The loading apparatus shall be constructed to allow the various increments of the 

load to be placed gradually, without causing vibration to the test pile.

When the approved method requires the use of reaction piles, the reaction piles 

shall be of the same type and diameter as the production piles. Reaction piles shall 

be surveyed and monitored for movement during the load test. Reaction piles driven 

at production pile locations that have a permanent upward displacement greater 

than 0.25 inch shall be redriven upon completion of the static load test. Timber or 

tapered piles installed in permanent locations shall not be used as reaction piles.

While performing the load test, the contractor shall provide safety equipment and 

employ adequate safety procedures. Adequate support for the load test plates, jack, 

and ancillary devices shall be provided to prevent them from falling in the event of a 

release of load due to hydraulic failure, test pile failure, or other cause.

The nominal geotechnical resistance or failure load of a pile statically tested in axial 

tension is defined by the pile head movement under load. The failure load is the pile 

head load which produces a measured upward movement of the pile head equal to:
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+∆= )15.0( f s Eq. 14-6 

Where: 

sf = pile head upward movement at failure (inches). 

Δ = elastic lengthening of the total pile length (inches). 

The maximum factored resistance under tension loading is the tension load test 

failure load determined using Equation 14-6 multiplied by the resistance factor, φup, 

for a static load test in axial tension from Table 14-2. 

Unless otherwise specified in the contract documents or by the Engineer, the static 

load test shall not be performed sooner than 5 days after the test pile or any anchor 

piles were driven.

On completion of the load testing, any test or anchor piling not a part of the finished 

structure shall be removed or cut off at least 1 foot below either the bottom of footing 

or the finished ground elevation, if not located within the footing area.

Commentary:

If the static load tests are to be performed by an independent firm retained by 

the Contractor and not by the Engineer, an additional specification section 

detailing the complete load test instrumentation monitoring requirements as 

well as the report submission requirements for the load test results and result 

interpretation must be added. A corresponding pay item must then be added 

to this specification for load test reporting. Alternatively, the report 

requirements can be described herein and then included as part of the static 

load test pay item.

It is essential that static load testing be performed after equilibrium conditions 

in the soil have re-established as discussed in the commentary of Section 

14.2.5.1 A. The time required to analyze and/or review the load test results 

should be specified so that the delay time is clearly identified. Static load 

testing is discussed in greater detail in Chapter 9 of this manual.

2. Dynamic Testing with Signal Matching: Dynamic measurements following 

procedures set forth in ASTM D4945 will be taken by the Engineer during the driving 

of piles designated as dynamic test piles. Signal matching analysis should be 

performed on representative data collected at the end of initial driving and at the 

beginning of all restrike events. Additional signal matching analysis may be 
315

performed as determined by the Engineer. The maximum factored resistance for 

tension loading can be taken as the signal matching determined shaft resistance 

value multiplied by the resistance factor, φup, for dynamic testing in Table 14-2. 

Signal matching results from restrike tests should be used in this evaluation.

Commentary: 

The commentary provided under the dynamic testing section for 

determination of the nominal axial compression resistance is applicable and 

should be reviewed for additional guidance.

C) Lateral Resistance:

The resistance and movement of piles subject to lateral loads shall be determined by 

the Engineer based on the following method.

1. Static Load Tests: Lateral load tests shall be performed by procedures set 

forth in ASTM D3966. Unless otherwise specified, the lateral load shall be applied 

incrementally following the standard loading procedure up to maximum applied load 

of 200% of the design lateral load. The lateral load shall be applied by a hydraulic 

jack acting between two piles, or between one pile and a reaction system. Testing 

equipment and measuring systems shall conform to ASTM D3966. A load cell and 

spherical bearing plate shall be used with the loading apparatus.

The Contractor shall submit to the Engineer for approval, detailed plans of the 

proposed loading apparatus prepared by a licensed professional engineer. The 

submittal shall also include calibrations for the hydraulic jack, load cell, and pressure 

gage conducted within 30 days of the load test. If requested by the Engineer, the 

jack, load cell, and pressure gage shall be recalibrated after the load test. The 

loading apparatus shall be constructed to allow the various increments of the load to 

be placed gradually, without causing vibration to the test pile.

Commentary: 

ASTM D3966 provides a standard and 7 optional lateral load test procedures. 

All of the procedures have different loading schedules. Therefore, the loading 

procedure should be clearly identified so that the time required to conduct the 

lateral test is a function of the procedure.
316

If the lateral load tests are to be performed by an independent firm retained 

by the Contractor and not by the Engineer, an additional specification section 

detailing the complete lateral load test instrumentation monitoring 

requirements as well as the report submission requirements for the load test 

results and result interpretation must be added. A corresponding pay item 

must then be added to this specification for lateral load test reporting. 

Alternatively, the report requirements can be described herein and then 

included as part of the lateral load test pay item.

Unless otherwise specified or directed by the Engineer, lateral pile deflection 

measurements versus depth shall be obtained for each lateral load increment 

applied during the test using a string of in-place inclinometers, a Shape Accel Array, 

or other approved measure.

While performing the lateral load test, the contractor shall provide safety equipment 

and employ adequate safety procedures. The load test plates, jack, and ancillary 

devices shall be restrained to limit movement in the event of a sudden release of 

load due to hydraulic failure, test pile failure, or other cause. 

14.2.6 SECTION X.06 PREPARATION AND DRIVING 

A) Site Work:

1. Excavation and Fill Placement: Where practical, the site grade in the 

immediate work area shall be excavated to the specified elevation before the piles 

are driven. Material forced up between and adjacent to driven piles within the cap 

area shall be removed to the required elevation prior to pile cap concrete placement.

At bridge abutments and other locations directly adjacent to fill placement, fill 

material shall be placed and the fill settlement complete to the magnitude specified 

by the foundation designer before driving piles.

Commentary: 

When approved by the Engineer, piles may be installed prior to embankment 

construction when minimal settlement or lateral displacement is expected in 

the soils beneath the embankment.
317

2. Predrilling for Driving: When required by the contract documents, the 

Contractor shall predrill holes of a size specified, at pile locations, and to the depths 

shown in the contract documents or approved in writing by the Engineer. After pile 

placement in the predrilled hole, the pile shall be driven with an impact hammer to 

the driving criteria specified by the Engineer. Any void space remaining around the 

pile in the predrilled zone after completion of driving shall be filled with sand or other 

approved material unless specifically prohibited or otherwise directed by contract 

documents. Material resulting from the drilled holes shall be disposed of as 

approved by the Engineer. The use of spuds shall not be permitted in lieu of 

predrilling, unless specifically approved in writing by the Engineer.

Commentary:

Except for end bearing piles, predrilling shall be stopped at least 5 feet above 

the plan pile toe elevation and the pile shall be driven with an impact hammer 

to a penetration resistance criteria specified by the Engineer. Where piles are 

to be end bearing on rock or hardpan, predrilling may be carried to the 

surface of the rock or hardpan.

If the Engineer determines that predrilling has disturbed the nominal resistance of 

previously installed piles, those piles that have been disturbed shall be restored to 

conditions meeting the requirements of this specification by redriving or by other 

methods acceptable to the Engineer. Redriving or other remedial measures shall be 

instituted after the predrilling operations in the area have been completed. The 

Contractor shall be responsible for the costs of any necessary remedial measures, 

unless the predrilling method was specifically included in the contract documents 

and properly executed by the Contractor.

Commentary:

Augering, wet rotary drilling, or other methods of predrilling shall be used only 

when approved by the Engineer or in the same manner as used for any 

indicator piles or load test piles. When permitted, such procedures shall be 

carried out in a manner which will not impair the nominal resistance already in 

place or the safety of existing adjacent structures.

3. Predrilling through Embankments: If required by contract documents, 

predrilled holes extending to natural ground shall be used where piles are to be 

driven through compacted fill or embankments greater than 5 feet in thickness. The 

predrilled hole shall have a diameter not more than the greatest dimension of the 
318

pile cross section plus 6 inches unless a different predrilled hole diameter is 

specified. Material resulting from the drilled holes shall be disposed of as approved 

by the Engineer.

B) Preparation of Piles for Driving:

1. Pile Heads: The heads of all piles shall be plane and perpendicular to the 

longitudinal axis of the pile before the helmet is attached. Precast concrete pile 

heads shall be flat, smooth, and perpendicular to the longitudinal axis to prevent 

eccentric impact from the helmet. Prestressing strands shall be cutoff below the 

surface of the end of the pile. The heads of all concrete piles shall be protected with 

a suitably thick pile cushion during driving as described in Section X.04 B) 3. For 

concrete and timber piles, the pile head shall be chamfered on all sides.

2. Collars and Bands: Timber pile heads shall be equipped with collars, 

bands, or other devices to protect against splitting and brooming of the pile head 

when timber piles are to be driven to a nominal resistance in excess of 200 kips or 

when driving conditions require them.

3. Pile Shoes and Closure Plates: Pile shoes and closure plates of the type 

and dimensions specified shall be provided and installed on piles when designated 

on the contract plans or specifications. Pile shoes for H-piles and open end pipe 

piles shall be fabricated from cast steel conforming to ASTM A148/A148M (Grade 

90-60). End closure plates for closed end pipe piles shall be made of ASTM 

A36/A36M steel or better. The closure plate diameter and thickness shall be as 

specified by the Engineer. Shoes for timber piles shall be steel and shall be 

fastened securely to the pile. Timber pile toes shall be carefully shaped to secure an 

even uniform bearing on the pile shoe.

Commentary:

H pile shoes composed of steel plates welded to the flanges and webs are 

not recommended because this reinforcement provides neither protection nor 

increased strength at the critical area of the flange to web connection. The 

designer should select and detail on the plans the proper pile shoe to suit the 

application. Additional information on pile shoes is presented in Chapter 16

of this manual.

4. Pile Marking for Prior to Driving: The Contractor shall mark the piles in 1 foot 

increments beginning at the pile toe and continuing to the pile head. The cumulative 
319

distance from the pile toe shall be marked on the pile at 5 foot intervals from the pile 

toe. These cumulative distances shall be noted just above the corresponding foot 

marker. Prior to driving, the Contractor shall, if necessary, add inch marks between 

the 1 foot markers over a 10 foot length of pile as directed by the Engineer.

C) Pile Driving:

1. Test, Probe, and Indicator Piles: Test, probe, or indicator piles (hereafter 

referred to collectively as test piles), shall be driven before other piles are ordered. 

These piles shall be driven at the locations shown on the plans and to the 

penetration depths, penetration resistances (blow count) or nominal resistances as 

directed by the Engineer. In general, the specified length of test piles will be slightly 

greater than the estimated length of production piles in order to provide for variation 

in soil conditions. The driving equipment used for driving test piles shall be the 

approved system that the Contractor proposes to use on the production piling. 

Approval of driving equipment shall follow the requirements of in Section X.02 of 

these Specifications. The Contractor shall excavate the ground at each test pile 

location to the elevation of the bottom of the footing before the test pile is driven.

Test piles which do not attain the nominal driving resistance before reaching a 

distance of 1 foot above the estimated pile toe elevation on the plans shall be 

allowed to "set up" for 12 to 24 hours, or as directed by the Engineer, before being 

redriven. A cold hammer shall not be used for redrive. The hammer shall be 

warmed up before driving begins by applying at least 20 blows to another pile. If the 

specified nominal driving resistance is not attained on redriving, the Engineer may 

direct the Contractor to drive a portion or all of the remaining test pile length and 

repeat the "set up" redrive procedure. Test piles that have not achieved the required 

nominal driving resistance during redrive shall be spliced, if necessary, and driven 

until the required nominal driving resistance is obtained or as directed by the 

Engineer.

A record of driving of the test pile will be prepared by the Engineer, including the 

number of hammer blows per 1 foot intervals over the entire driven length, the asdriven length of the test pile, cutoff elevation, penetration in ground, and any other 

pertinent information. Near the end of initial driving, the Engineer may record the 

number of hammer blows per 1 inch of pile movement. If a redrive is necessary, the 

Engineer will record the number of hammer blows per 1 inch of pile movement for 

the first 1 foot of redrive. The Contractor shall not order piling to be used in the 

permanent structure until test pile data has been reviewed and pile order lengths are 

determined by the Engineer. The Engineer will provide the pile order list within 7 
320

calendar days after completion of all test pile driving specified in the contract 

documents.

Commentary:

Test piles are recommended on projects where: 1) large quantities or long 

lengths of friction piling are estimated, even if load tests are to be used at 

adjacent footings; 2) large nominal soil resistance is expected in relation to 

the factored load and, 3) concrete piles are used. 

The pile order lengths based on the test pile results should consider the 

anticipated variation in subsurface conditions. This is a particularly important 

consideration for concrete or timber piles which are both problematic if 

ordered shorter than required.

2. Production Piles: Production piles shall be driven by the Contractor to either 

the required nominal resistance, or the required nominal resistance and minimum 

pile toe elevation (if specified), or to a designated pile toe elevation. The pile 

penetration resistance (blow count) and the associated pertinent hammer 

performance observation for the hammer type (hammer stroke, impact velocity, 

hammer blow rate) should be documented during all driving sequences. Pile 

penetration resistance and the corresponding hammer performance documentation 

are required to evaluate the nominal resistance of impact driven piles.

Jetting or other methods shall not be used to facilitate pile penetration unless 

specifically permitted in the contract plans or in writing by the Engineer. The 

nominal resistance of jetted piles shall be based on the penetration resistances 

recorded during impact driving after the jet pipes have been removed. Jetted piles 

not attaining the nominal resistance at the ordered length shall be spliced, as 

required, at the Contractor's cost, and driven with an impact hammer until the 

nominal resistance is achieved using the approved driving criteria.

The nominal resistance of piles driven with followers shall be considered acceptable 

when the follower driven piles satisfy the driving criteria established by wave 

equation analysis and the criteria is either substantiated or modified based on a 

static load test or dynamic testing with signal matching.

The nominal resistance of piles driven with vibratory hammers shall be based on the 

penetration resistance recorded during impact driving after the vibratory equipment 

has been removed from the first pile in each group of ten piles. Vibratory installed 
321

piles not attaining the nominal resistance at the estimated pile toe elevation shall be 

spliced, at the Contractor's cost, and driven with an impact hammer until the nominal 

resistance is achieved. When the nominal resistance is attained, the remaining 9 

piles shall be installed to similar pile toe elevation with similar vibratory hammer 

power consumption and rate of penetration as the first pile.

Approval of a pile hammer relative to material driving stress levels shall not relieve 

the Contractor of responsibility for piles damaged because of misalignment of the 

leads, deterioration of cushion materials, failure of splices, malfunctioning of the pile 

hammer, or other improper construction methods. Piles damaged for such reasons 

shall be rejected and replaced at the Contractor's expense when the Engineer 

determines that the damage impairs the strength of the pile.

3. Driving Stress: Compression and tension stresses occurring in the pile 

material during driving shall not exceed the maximum stress levels defined in 

Section X.02 B) unless otherwise specified in the contract documents or by the 

Engineer.

4. Installation Sequence: The order of placing and final driving of individual 

piles within pile groups shall be either starting from the center of the group and 

proceeding outwards in both directions or starting at the outside row and proceeding 

progressively across the group.

5. Pile Location, Alignment, and Cutoff Tolerance: Piles shall be driven with 

a variation of not more than 0.25 inches/foot (1:50) from the vertical or not more than 

0.5 inches /foot (1:25) from the batter shown in the contract documents. Piles for 

trestle bents shall also be driven so that the bent cap may be placed in its proper 

location without adversely affecting the resistance of the piles.

The pile head location after driving shall be within 6 inches of plan location for all 

piles capped below final grade, and shall be within 3 inches of plan location for bents 

supported by piles.

No pile shall be closer than 4 inches from any edge of the pile cap. Any increase in 

pile cap dimensions or added reinforcing steel caused by out-of-tolerance or out-ofposition piles shall be at the Contractor’s expense.

The final cutoff elevation of the pile head shall be no more than 1.5 inches above or 

more than 4 inches below the cutoff elevation shown in the plans. In addition, the 

pile shall have a minimum embedment into the pile cap of at least 8 inches.
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If the location and/or tolerances specified in the preceding paragraphs are 

exceeded, the extent of overloading shall be evaluated by the Engineer. The cost of 

redesign shall be at the Contractor’s expense.

Commentary: 

Conditions exist, such as soft overburden soils directly overlying a sloping 

bedrock, where final pile location and/or alignment may be beyond the 

contractor's control. These cases should be identified during the design stage 

with specifications tailored to meet the site and project requirements. Tight 

tolerances of 3 inches or less are not practical.

6. Heaved Piles: Level readings to check for pile heave after driving shall be 

made by the Engineer at the start of pile driving operations and shall continue 

periodically until the Engineer determines that such checking is no longer required. 

If pile heave is observed, accurate level readings referenced to a fixed datum shall 

be undertaken by the Engineer on all piles immediately after installation and 

periodically thereafter as adjacent piles are driven to determine the magnitude of the 

pile heave. Piles that derive their nominal resistance predominant through end 

bearing shall be redriven if more than 0.25 inch of heave is measured. Piles that 

achieve their nominal resistance primarily through shaft resistance shall be redriven 

if more than 1.5 inches of heave is measured. If pile heave is detected on any piles 

that have been concrete filled, the piles shall be redriven to original position after the 

concrete has obtained sufficient strength. Redriving shall be done using a hammerpile cushion system satisfactory to the Engineer. The Contractor shall be paid for all 

work performed in conjunction with redriving piles because of pile heave provided 

the initial driving was done in accordance with the specified installation sequence 

and approved pile installation plan.

7. Obstructions: If piles encounter unforeseeable, isolated obstructions, the 

Contractor shall be paid for the cost of obstruction removal and for all remedial 

design or construction measures caused by the obstruction. Obstruction removal is 

only practical when obstructions are located near the ground surface.

8. Practical and Absolute Refusal: Practical refusal is defined as a pile 

penetration resistance (blow count) of 10 blows per inch for a maximum of 3 

consecutive inches of pile penetration. Absolute refusal is defined as 20 blows for 

one inch or less of pile penetration. Driving should terminate immediately if either of 

these criteria are achieved. In the case of hard rock, an absolute refusal criterion of 

5 blows per 1⁄4 inch or 10 blows per 1⁄2 inch may be preferred to reduce the risk of pile 
323

toe or driving equipment damage. The practical and absolute refusal definitions are 

based on the approved hammer system operating at its maximum fuel or stroke 

setting unless approval was established on hammer operation at a reduced fuel or 

stroke setting. When the required pile penetration depth cannot be achieved by 

driving without exceeding practical or absolute refusal with the approved hammer, 

use of other penetration aids such as predrilling or jetting should be evaluated. 

Commentary:

Clear definitions for practical and absolute refusal are problematic. Factors 

such as the site soil profile, the characteristics of the bearing layer, pile type, 

hammer type, and hammer manufacturer limitations to prevent hammer 

damage all influence what is considered an acceptable refusal criteria. Many 

hammer manufacturers state that the hammer warranty is voided when pile 

penetration resistance (blow count) exceed 10 blows per inch for 6 to 12 

consecutive inches of driving or 20 blows per inch for more than 1 inch.

When driving is easy until final driving, a high penetration resistance may be 

satisfactory. However, penetration resistances greater than 10 blows per inch 

should be used with care, particularly for concrete or timber piles. Extended 

driving at a penetration resistance greater than 10 blows per inch with a 

properly operating hammer should be avoided.

9. Pile Splices: Full length piles shall always be used wherever practical. 

Where splices are unavoidable for steel or concrete piles, their number, locations 

and details shall be subject to approval of the Engineer. In no case shall timber piles 

be spliced.

When splicing of steel piles is permitted, the method of splicing piles shall be in 

accordance with ANSI/AWS D1.1 or as approved by the Engineer. Either shielded 

arc or submerged arc welding should be used when splicing steel piles. Only 

certified welders shall perform welding. Mechanical splices that are not welded may 

be used only for compression piles.

Where splicing concrete piles is permitted, the concrete pile splice details shall 

conform to contract documents or as approved by the Engineer. Mechanical splices 

may be used if they satisfy all compression, tension, and bending resistance 

requirements of the design.
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10. Pile Cutoff: The pile head of all permanent piles and pile casings shall be 

cutoff to a true plane at the required elevation and anchored to the structure as 

shown on the contract documents. All cutoff lengths shall become the property of 

the Contractor, and shall be removed from the site and disposed of properly.

For treated timber piles, a liberal application of copper naphthenate shall be given to 

the cut area immediately after cutoff. The copper naphthenate solution shall have a 

minimum of 2 percent copper metal and should be applied until visible evidence of 

further penetration has ceased.

Treated timber piles for marine applications exposed to weather shall be capped 

with a permanently fixed coating such as epoxy or with conical or other caps 

attached to the piles.

Commentary:

Additional structural details for timber, steel, concrete and cast in place piles 

should be included by each agency in this driven pile specification, either 

directly or by reference to appropriate sections of the individual agency's 

standard specification. Typical items include: timber pile butt treatment and 

preservative treatment; precast concrete pile reinforcement, forming, casting, 

curing, and handling; steel pile field painting; cast in place pile details for shell

piles, interior reinforcement and concrete.

11. Unsatisfactory Piles: The methods used in transportation, handling, and 

driving piles shall not subject the piles to excessive stresses or abuse producing 

cracking, crushing or spalling of concrete piles; splitting, splintering, or brooming of 

the timber piles; or deformation of steel piles. A concrete pile will be considered 

defective if a visible crack, or cracks, appears around the entire periphery of the pile, 

or if any defect is observed which, as determined by the Engineer, affects the 

strength or service life of the pile. Misaligned piles shall not be forced into proper 

position. Any pile damaged during driving by reason of internal defects, or by 

improper driving, or by defective splicing, or driven out of its proper location, or 

driven below the designated cutoff elevation, shall be corrected by the Contractor, 

without added compensation, by a method approved by the Engineer.

Commentary:

Defective piles can often be remediated by: 
325

• withdrawing the defective pile and replacing it with a new and, when 

necessary, longer pile. In removing piles, jets may be used in 

conjunction with jacks or other devices for pulling in an effort to remove 

the whole pile, or

• driving a second pile adjacent to the defective pile.

Piles driven past their specified top of pile elevation can:

• be spliced or built up as otherwise provided herein, or

• have a sufficient portion of the footing extended down to properly 

embed the overdriven pile.

Piles driven out of location can be remediated by:

• driving one or more replacement piles adjacent to the out of position 

piles, or

• extending the footing laterally to incorporate the out of location pile, or

• redesigning and adding more reinforcing steel to the pile cap.

14.2.7 SECTION X.07 METHOD OF MEASUREMENT

A) Mobilization of Pile Driving Equipment:

Payment for mobilization will be made at the lump sum price bid for this item as 

follows: Seventy five percent (75 percent) of the amount bid will be paid when the 

equipment for driving piles is mobilized, assembled, and driving of satisfactory piles 

has commenced. The remaining 25 percent will be paid when the work of driving 

piles is completed. The lump sum price bid shall include the cost of furnishing all 

labor, materials and equipment necessary for transporting, erecting, maintaining, 

replacing any ordered equipment, dismantling and removing of the entire pile driving 

equipment.

The cost of all labor, including the manipulation of the pile driving equipment and 

materials in connection with driving piles, shall be included in the unit price bid for 

the piles to be driven.
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B) Piles Furnished:

The quantity of piles furnished will be measured for payment using the number of 

piles furnished (per linear foot or each) in accordance with the contract documents

and shall include full compensation for all costs involved with furnishing and 

delivering piles to the project site in the unit price bid for furnished piles.

When pile build-ups or extensions are necessary, the extension length, approved by 

the Engineer, will be included in the total length of piling furnished.

No allowance will be made for that length of piles furnished by the Contractor to 

replace piles which were previously accepted by the Engineer, but were

subsequently damaged by the Contractor prior to completion of the contract.

C) Piles Driven:

The quantity of piles driven will be measured for payment using the number of piles 

driven (per linear foot or each) in accordance with the contract documents and shall 

include full compensation of all costs involved in the actual driving of the piles as 

well as all related pile costs for which compensation is not identified as a specific 

pay item. These related pile costs shall include furnishing all labor, equipment, and 

materials necessary to install and complete the pile. All costs shall be included in 

the unit price bid for the piles driven.

D) Test Piles Furnished:

The quantity of test piles furnished will be measured for payment using the number 

test piles furnished (per linear foot or each) in accordance with the contract 

documents and shall include full compensation of all costs involved with furnishing 

and delivering test piles to the project site in the unit price bid for furnished test piles.

E) Test Piles Driven:

The quantity of test piles driven will be measured for payment using the number of 

test piles driven (per linear foot or each) in accordance with the contract documents

and shall include full compensation of all costs involved in the actual driving of test 

piles as well as all related test pile costs for which compensation is not identified as 

a specific pay item. These related test pile costs shall include furnishing all labor, 

equipment, and materials necessary to install and complete the test pile. All costs 

shall be included in the unit price bid for the test piles driven.
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F) Static Pile Load Test:

The quantity of static load tests for payment will be measured by the number of load 

tests completed and accepted. This pay item shall include all labor, equipment, and 

materials needed to construct and perform the static pile load test in accordance 

with contract documents.

Reaction and test piling which are not a part of the permanent structure will be 

included in the unit price bid for each load test. Reaction and test piling, which are a 

part of the permanent structure, will be paid for under the appropriate pay item.

Static load tests performed at the option of the Contractor will not be included in the

quantity measured for payment.

G) Dynamic Pile Load Test:

The quantity of dynamic pile tests for payment will be measured by the number of 

dynamic pile tests completed and accepted in accordance with the contract 

documents. The pay item for dynamic pile test (during driving), or dynamic pile test

(during restrike) shall include full compensation for furnishing all labor, equipment, 

and materials necessary to perform the dynamic pile test as specified in the contract 

documents. If the dynamic test is performed at a test pile location (non-production 

location), the unit price for test piles furnished and test piles driven will be paid in 

addition to the unit price for the dynamic pile test (during driving), or dynamic pile 

test (during restrike). If the dynamic test is performed at a production pile location, 

the unit price for piles furnished and piles driven will be paid in addition to the unit 

price for the dynamic pile test (during driving), or dynamic pile test (during restrike). 

If an unspecified dynamic pile test requires substantial repositioning, delay, or 

downtime of the pile driving rig (such as may occur for a second longer term restrike 

on a test pile conducted at the owner’s request), additional compensation shall be 

paid at the unit price per hour for the out of sequence move, delay, or downtime in 

addition to the applicable unit bid prices for dynamic pile test, test piles, or piles.

H) Rapid Pile Load Test:

The quantity of rapid load tests for payment will be measured by the number of load 

tests completed and accepted. This pay item shall include all labor, equipment, and 

materials needed to prepare the site for the test, construct and perform the test per 

the contract documents.
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I) Lateral Load Test:

The quantity of lateral load tests for payment will be measured by the number of 

lateral load tests completed and accepted. When contract documents stipulate that 

two piles are to be laterally load tested by pushing apart or pulling together two piles, 

the lateral load test pay quantity shall comprise the lateral load testing of both piles. 

This pay item shall include all labor, equipment, and materials needed to prepare the 

site for the test, construct and perform the lateral load test per the contract 

documents.

J) Splices:

The quantity of splices measured for payment shall be only those splices actually 

made that were required to drive the piles in excess of the plan estimated and 

approved pile lengths as accepted for payment by the Engineer. The unit price bid 

per splice shall comprise full compensation for procurement, delivery, and 

attachment of the splice including all labor, equipment, and ancillary materials.

K) Pile Shoes:

The quantity of pile shoes measured for payment shall be those shoes actually 

installed on piles and accepted for payment by the Engineer. The unit price bid per 

pile shoe shall comprise full compensation for procurement, delivery, and 

attachment of the shoes including all labor, equipment, and ancillary materials.

Commentary

Pile shoes can be alternatively be included in the furnished pile price if clearly 

stated on the plans and in the contract documents.

L) Predrilling:

The quantity of predrilling measured for payment shall be taken to include full 

compensation for providing all labor, equipment, and materials necessary to perform 

the predrilling work in accordance with the contract documents. 
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M) Jetting

The quantity of jetting measured for payment shall be taken to include full 

compensation for providing all labor, equipment, and materials necessary to perform 

jetting work in accordance with the contract documents. 

N) Pile Cutoff

The quantity of pile cutoffs measured for payment shall be taken to include full 

compensation for providing all labor and equipment needed to provide a plane and 

level pile head surface at the required cutoff elevation as specified by the contract 

documents. The contract unit price also includes full compensation for proper 

disposal of the cutoff material.

O) Spudding

The quantity of spudding measured for payment shall be taken to include full 

compensation for providing all labor, equipment, and materials necessary to perform 

the spudding work in accordance with the contract documents. 

P) Delays, Downtime, Rig Moves

The quantity of time measured for payment for delays, downtime, or out of sequence 

pile driving rig moves caused by the owner, agents, or subcontractors not otherwise 

compensated in contract pay items shall be recorded for each claimed occurrence 

and shall be taken to include full compensation for all labor and equipment.

14.2.8 SECTION X.08 BASIS OF PAYMENT

The accepted quantities, determined as described above, will be paid for at the 

respective contract document price per unit of measurement for each of the general 

pay items listed below. Payment will be made for each size and type of pile shown 

in the contract documents. Prices and payment will be full compensation for the 

work prescribed by the contract documents. Pay item and pay units are described 

below in Table 14-5. 
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Table 14-5 Basis of Payment

Pay Item Pay Unit

Mobilization and Demobilization Lump sum

Piles Furnished Linear foot or each

Piles Driven Linear foot or each

Test Pile Furnished Linear foot or each

Test Pile Driven Linear foot or each

Static Load Test Each

Dynamic Pile Test (during driving) Each

Dynamic Pile Test (during restrike) Each

Rapid Load Test Each

Lateral Load Test Each

Pile Splices Each

Pile Shoes Each

Predrilling Linear foot or Each

Jetting Linear foot or Each

Pile Cut-off (over 5 ft lengths only) Each

Spudding (Punching) Per hour

Delays, Downtime, or Out-of-Sequence Moves Per hour

Commentary: 

The pile payment items in Table 14-5 have been chosen to separate the 

major fixed costs from the variable costs. Many agencies oversimplify pile 

payment by including all costs associated with the driving operation in the 

price per foot of pile installed. Contractors bidding such "simple" items need 

to break down the total cost of the mobilization, splices, shoes, etc., to a price 

per linear foot based on the total estimated quantity. If that quantity 

underruns, the contractor does not recover the full cost of mobilization, 

splices, shoes, etc. If that quantity overruns, the agency pays an unfair price 

for the overrun quantity. The use of separate items for operations of major 

fixed cost such as mobilization can substantially mitigate the inequitable 

impact of length variations. Similarly, the ordered pile length is the agency's 

responsibility. Separate payment for furnishing piles and driving piles 

compensates the contractor for actual materials used and installation costs, 

even when modest overruns or underruns occur.
331

REFERENCES

American Association of State Highway and Transportation Officials (AASHTO). 

(2010). AASHTO LRFD Bridge Construction Specifications, US Customary 

Units, Third Edition, with 2011, 2012, 2013, 2014, and 2015 Interim 

Revisions. American Association of State Highway and Transportation 

Officials, Washington, D.C., 680 p.

American Association of State Highway and Transportation Officials (AASHTO). 

(2014). AASHTO LRFD Bridge Design Specifications, US Customary Units, 

Seventh Edition, with 2015 Interim Revisions. American Association of State 

Highway and Transportation Officials, Washington, D.C., 1960 p.

ASTM D1143-07. (2014). Standard Test Methods for Deep Foundations Under 

Static Axial Compressive Load. Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 15 p.

ASTM D3689-07. (2014). Standard Test Methods for Deep Foundations Under 

Static Axial Tensile Load. Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 13 p.

ASTM D4945-12. (2014). Standard Test Method for High-Strain Dynamic Testing of 

Piles. Book of ASTM Standards, Vol. 4.08, ASTM International, West 

Conshohocken, PA, 9 p.

Deep Foundations Institute (DFI). (1995). A Pile Inspector's Guide to Hammers, 

Second Edition, Deep Foundations Institute, Englewood Cliffs, NJ, pp. 71.

Deep Foundations Institute (DFI). (1997). Inspectors Manual for Driven Pile 

Foundations, Second Edition, Deep Foundations Institute, Englewood Cliffs, 

NJ, pp. 69.

Geotechnical Guideline 13 (1985). Geotechnical Engineering Notebook, U.S. 

Department of Transportation. Federal Highway Administration, Washington, 

D.C., 37 p.
332

Rausche, F., Likins, G.E., Goble, G.G., Hussien, M. (1986). The Performance of Pile 

Driving Systems; Inspector's Manual. U.S. Department of Transportation, 

Federal Highway Administration, Washington, D.C., 92 p.
333

CHAPTER 15

PILE DRIVING EQUIPMENT

The task of successfully installing piles involves selecting the most cost effective 

equipment to drive each pile to its specified depth without damage in the least 

amount of time. The pile driving system is also used as a measuring instrument to 

evaluate a pile’s nominal geotechnical resistance. Therefore, the challenge to both 

the engineer and the pile contractor becomes one of knowing, or learning about, the 

most suitable equipment for a given set of site conditions, and then confirming that 

the driving system is operating properly.

The crane, leads, hammer, and helmet are the primary components of any driving 

system. Followers and equipment for jetting, predrilling, and spudding may be 

permitted under certain circumstances. This chapter presents a basic description of 

each component of a driving system. For additional guidance, readers are referred 

to pile driving equipment manufacturers and suppliers. Unless otherwise noted, 

photographs are provided courtesy of GRL Engineers, Inc.

15.1 CRANES 

The most common crane used for pile driving has historically been the crawler 

crane. Crawler cranes have very good mobility for most site conditions, good 

stability due to their wide base, and typically require minimal effort when walked and 

repositioned on the job site. A hydraulic power pack unit or air compressor may be 

mounted as a crane counterweight to facilitate pile driving operations. Crawler 

cranes also have a 360 degree hoisting radius. Crawler cranes must be trucked to 

the job site in pieces and assembled by another crane. Firm ground conditions are 

required for movement and operation. In soft ground conditions, the use of crane 

mats may be required. A photograph of a typical crawler crane set up for pile driving 

is presented in Figure 15-1. 

Truck or mobile cranes are also used as pile driving rigs. Mobile cranes have good 

mobility on many job sites and can travel to the job under their own power. Setup 

and breakdown are relatively simple. However, mobile cranes cannot walk with 

heavy loads. Therefore, mobile crane movement between substructure locations 
334

Figure 15-1 Crawler crane pile driving rig with lattice boom and swinging leads.

usually requires laying down and un-attaching the hammer and leads, moving the 

crane, and then re-attaching the hammer and lead system. Outriggers are used on 

mobile cranes for stability and leveling. The power pack unit or air compressor 

cannot be crane mounted. Mobile cranes also have restricted side lifting 

capabilities. A photograph of a mobile crane set up for pile driving is presented in 

Figure 15-2. 
335

Figure 15-2 Truck crane pile driving rig with lattice boom and swinging leads.

Rough terrain cranes can also be used for pile driving. They can be set up fast and 

easily on a job site, they can be relocated and releveled relatively quickly, and can 

walk in some cases with loads over the front. When on location, outriggers are used 

for stability and leveling. Rough terrain cranes have greater mobility than truck 

cranes in difficult terrain. However, they have very restricted side lifting capabilities

and a hammer power source cannot be crane mounted. A photograph of a rough 

terrain crane set up for pile driving is presented in Figure 15-3. 
336

Figure 15-3 Rough terrain pile driving rig with telescoping boom and offshore leads.

15.2 DEDICATED AND UNIVERSAL RIGS

In recent years, use of dedicated and universal rigs has increased on pile driving 

projects. These specialty pile driving rigs or universal rigs include an attached 

telescoping leader that can be quickly transitioned from their horizontal transport 

position to the vertical driving position using hydraulically operated support arms. 

Hence, system setup is relatively fast without the need of additional heavy 

equipment. These systems include winch lines that can be used to offload piles as 
337

well as loft piles into position beneath the hammer prior to pile driving. A dedicated 

pile driving system equipped with a hydraulic pile hammer is presented in Figure 

15-4, while a universal rig setup for driving piles with a diesel pile hammer is shown 

in Figure 15-5. 

Figure 15-4 Dedicated pile driving system.
338

Figure 15-5 Universal rig setup for pile driving (courtesy Bauer-Pileco).

15.3 LEADS

Lead systems are used with crawler, truck, and rough terrain pile driving rigs. The 

function of a set of leads is to maintain alignment of the hammer-pile system so that 

a truly concentric blow is delivered to the pile for each impact. Typical lead types 

are illustrated in Figure 15-6. The most common lead type is the box lead. Leads 

can be configured for use with the pile driving rig as a swinging lead, fixed lead, 

semi-fixed or vertical travel lead, or as an offshore lead. The most widely used lead 

configuration is a swinging lead depicted in Figure 15-7. Swinging leads are widely 

used because of their simplicity, lightness and low cost. Leads can be hung from
339

Figure 15-6 Typical lead types (after DFI Publication 1981).
340

Figure 15-7 Swinging lead systems (after DFI Publication 1981).
341

the boom with hanger straps as illustrated in Figure 15-7(a) with the hammer held by 

a crane line. The most common arrangement however, is shown in Figure 15-7(b) 

where the leads and hammer are held by separate crane lines. Swinging leads are 

free to rotate sufficiently to align the hammer and the head of the pile without precise 

alignment of the crane with the pile head. Because the weight of the leads is low, 

this type of lead generally permits the largest crane operating radius, providing more 

site coverage from one crane position.

Photographs of swinging lead arrangements on pile driving rigs were previously 

presented in Figure 15-1 and 15-2. A stabbing guide is located at the bottom of a 

swinging lead system as illustrated in Figure 15-8. Pile location and alignment are 

controlled by positioning the lead system over the pile location, lifting the leads, and 

then dropping the leads so that the stabbing guide at the bottom of the leads 

penetrates into the ground. The crane boom is then manipulated until the specified 

verticality or batter angle is achieved.

Pile driving specifications have historically penalized or prohibited swinging leads. 

This general attitude is not justified based on currently available equipment. There 

are many cases where swinging leads are more desirable than fixed leads. For 

example, swinging leads are preferable for pile installation in excavations or over 

water. As noted earlier, the function of a lead is to hold the pile in good alignment 

with the driving system in order to prevent pile damage, and to hold the pile in its 

proper position for driving. If a swinging lead is long enough so that the bottom is 

firmly embedded in the ground, and if the bottom of the lead is equipped with a gate 

then bottom alignment of the pile will be maintained (Figure 15-9). In this situation, if 

the pile begins to move out of position during driving, it must move the bottom of the 

lead with it. Swinging leads should be of sufficient length so that the free line 

between the boom tip and the top of the leads is short, thus holding the top of the 

lead in good alignment. When batter piles are driven, pile alignment is more difficult 

to set with swinging leads. This problem is accentuated for diesel hammers since 

the hammer starting operation will tend to pull the pile out of line.

Fixed leads are connected to the boom point and to the crane frame using a spotter 

or brace that runs from the bottom of the leads to the crane. A schematic of a typical 

fixed lead system is depicted in Figure 15-10(a) and a photograph is presented in 

Figure 15-11. A fixed lead system attempts to hold the pile in true alignment while 

driving but may require more time for rig repositioning prior to driving.
342

Figure 15-8 Stabbing guide at bottom of box leads.

Figure 15-9 Pile gate with latch for use on truss lead (courtesy Berminghammer).
343

Figure 15-10 Fixed lead systems (after DFI Publication 1981).
344

Figure 15-11 Fixed lead system with air compressor mounted as counterweight 

(courtesy FHWA Demo 66).

As an alternative to a fixed lead system, a semi-fixed or vertical travel lead may be 

used. A semi-fixed lead, as shown in Figure 15-12, allows vertical lead movement at 

the lead connection points to the boom and brace which the standard fixed lead 

system does not. 
345

Figure 15-12 Vertical travel lead system (courtesy Berminghammer).

Figure 15-13(a) illustrates that a fixed lead is limited to plumb piles or batter piles in 

line with the leads and crane boom. To drive side batter piles, a moonbeam must be 

attached at the end of the spotter or brace as depicted in Figure 15-13(b). 

Offshore leads are similar to swinging leads in that they are free to rotate sufficiently 

to align the hammer and pile head, however they do not require precise alignment of 

the crane with the pile head, and they generally consist of only a short lead section 

with a pile guide at the base. An offshore lead schematic is depicted in Figure 

15-14. The short lead section is intended only to hold and axially align the hammer 

with the pile head, and does not provide support for batter or full pile alignment 

during the driving process. When offshore leads are used, a template is constructed 

to hold the pile in position during driving. Section 15.3 provides a further discussion 

on templates. A photograph of an offshore lead with guide is shown in Figure 15-15

and of an offshore lead system in use with a rough terrain crane in Figure 15-3. 
346

Figure 15-13 Lead configuration for batter piles (after DFI publication 1981).
347

Figure 15-14 Typical offshore lead configuration.

Figure 15-15 Offshore lead and pile guide (left) with pipe pile helmet (right)

(courtesy Berminghammer).

Regardless of lead type chosen, the pile must be kept in good alignment with the 

hammer to avoid eccentric impacts which can cause local stress concentrations and 

pile damage. The hammer and helmet, centered in the leads and on the pile head, 

keep the pile head in alignment. A pile gate at the bottom of the leads should be 

used to keep the lower portion of the pile centered in the leads.
348

15.4 TEMPLATES

Templates are required to hold piles in proper position and alignment when an 

offshore or swinging lead system is used over water or excavations. The top of the 

template should be located within 5 feet of the pile cutoff elevation or the water 

elevation, whichever is lower. The preferred elevation of the template is at or below 

the pile cutoff elevation so that final driving can occur without stopping for template 

removal. A photograph of a typical template is presented in Figure 15-16. When 

templates include batter piles, it must be remembered that the correct location for 

the batter piles in the template arrangement will vary depending upon the template 

elevation relative to the pile cutoff elevation. For example, consider a template 

located 5 feet above pile cutoff elevation. If the plan pile locations at cutoff are used 

at the template elevation, a 1H:4V batter pile would be 15 inches out of location at 

the pile cutoff elevation. This problem is illustrated in Figure 15-17. Template 

construction should also allow the pile to pass freely through the template without 

binding. Templates with rollers are preferable, particularly for batter piles.

Figure 15-16 Template system.
349

Figure 15-17 Template elevation effects on batter piles (after Passe 1994).

15.5 HELMETS

Figure 15-18 shows the components of a typical helmet and the nomenclature used 

for associated components. The helmet configuration and size used depends upon 

the lead type, pile type and the type of hammer used for driving. Helmets may be a 

one piece unit manufactured for driving a specific pile type and size, or may consist 

of two pieces consisting of a base helmet with an insert to accommodate various pile 

types and sizes. Details on the proper helmet for a particular hammer can be 

obtained from hammer manufacturers, suppliers and contractors. To avoid the 

transmission of torsion or bending forces, the helmet or insert should fit loosely, but 

not so loosely as to prevent the proper alignment of hammer and pile. Helmets or 

inserts should be approximately 0.1 to 0.2 inches larger than the pile diameter. 

Proper hammer-pile alignment is particularly critical for precast concrete piles.

In Figure 15-19 a photograph of a helmet that cracked during driving provides a 

unique cross sectional view illustrating a typical striker plate, hammer cushion 

(consisting of two aluminum plates and one micarta plate), and helmet configuration. 

Photographs of one piece helmets for a concrete pile and pipe pile are presented in 

Figure 15-20(a) and 15-20(b), respectively. 
350

Figure 15-18 Helmet components (after DFI Publication 1981).

Figure 15-19 Cross section of cracked helmet with striker plate and aluminum and 

micarta hammer cushion (courtesy of WKG2

).
351

Figure 15-20 One piece helmets for (a) concrete pile and (b) pipe pile.

A two piece system consisting of a base or primary helmet and an insert can also be 

used to adapt a base helmet for driving different pile types. A base helmet with an 

insert for driving H-piles is presented in Figure 15-21(a) and a photograph showing 

several pipe pile inserts to accommodate different pipe pile diameters and wall 

thickness on the same project is presented in Figure 15-21(b).

Figure 15-21 Inserts for (a) H-pile and (b) varying pipe pile diameter and wall 

thickness. 
352

15.6 HAMMER CUSHIONS

Most hammers use a hammer cushion between the hammer and the helmet to 

relieve the impact shock, thus protecting the pile hammer. However, some hammer 

models exist that do not require a hammer cushion, or utilize a direct drive option 

where the hammer cushion is replaced by a steel striker plate. Ineffective hammer 

cushions can cause damage to hammer striking parts, anvil, helmet or pile. 

Hammer cushion materials are usually proprietary man-made materials such as 

micarta, nylon, urethane or other durable polymers. Over time, hammer cushion 

materials become compressed and stiffen as additional hammer impacts are 

applied. Therefore, hammer cushions eventually become ineffective, or may result 

in significant reduction in transferred energy or increased bending stress. 

Proprietary hammer cushions may last for up to 200 hours of driving.

In the past, a commonly used hammer cushion was made of hardwood (one piece), 

approximately 6 inches thick, with the wood grain parallel to the pile axis. This type 

of cushioning has the disadvantage of quickly becoming crushed and burned as well 

as having variable elastic properties during driving. With the widespread availability 

of hammer cushions from durable man-made materials, hammer cushions consisting 

of hardwood, small pieces of wood, wire rope, or other highly elastic material should 

not be permitted. Cushion materials containing asbestos are not acceptable 

because of health hazards. All of these hammer cushion materials are prohibited by 

AASHTO LRFD Bridge Construction Specifications.

Proprietary man-made hammer cushion materials have better energy transmission 

characteristics than a hardwood block, maintain more nearly constant elastic 

properties, and have a relatively long life. Their use results in more consistent 

transmission of hammer energy to the pile and more uniform driving. Since 

proprietary hammer cushion materials last up to 200 hours, it is often sufficient on 

smaller projects to inspect the cushion material only once before the start of pile 

driving operation. Periodic inspections of hammer cushion wear and thickness 

should be performed on larger projects. Many hammers require a specific cushion 

thickness for proper hammer timing. In these hammers, improper cushion thickness 

will result in poor hammer performance. Some man-made hammer cushions are 

laminated sandwiches of aluminum and another material such as micarta. The 

aluminum is used to transfer the heat generated during impact out of the cushion, 

thus prolonging its useful life. Some common proprietary hammer cushion materials 

are shown in Figure 15-22 and include Conbest and aluminum (top left), Blue Nylon 

(right), and aluminum and Micarta (lower left).
353

Figure 15-22 Typical hammer cushion materials. 

15.7 PILE CUSHIONS

To avoid damage to the head of a concrete pile as a result of direct impact from the 

helmet, a pile cushion should be placed between the helmet and the pile head. 

Typical pile cushions are made of compressible material such as plywood, 

hardwood, plywood and hardwood composites or other man made materials. Wood 

pile cushions should have a minimum thickness of 4 inches. Pile cushions should 

be checked periodically for damage and replaced before excessive compression or 

charring takes place. After replacing a cushion during driving, the blow count from 

the first 100 blows should not be used for pile acceptance as the cushion is still 

rapidly absorbing energy. The blow count will only be reliable after 100 blows of full 

energy application. The total number of blows which can be applied to a wood 

cushion is generally between 1000 and 2000. For wood pile cushions, it is 

recommended that a new, dry cushion be used for each pile. Old or water soaked 

cushions do not have good energy transfer, and will often deteriorate quickly. A 
354

photograph of plywood pile cushions is presented in Figure 15-23. An unused 

cushion is shown on the right, while a used and compressed cushion is shown on 

the left.

Figure 15-23 Plywood pile cushions: (left) used and (right) new. 
355

15.8 HAMMERS

Pile hammers can be categorized in three main types: impact hammers, vibratory 

ahammers, and resonant hammers. Impact hammer are the primary hammer type 

used to install foundation piles. There are numerous types of impact hammers 

having variations in the types of power source, configurations, and rated energies. 

Figure 15-24 shows a classification of hammers based on motivation and 

configuration factors. Table 15-1 presents characteristics and uses of several types 

of hammers. A detailed discussion of various types of hammers follows later in this 

chapter. Discussion on the key inspection issues associated with each hammer type 

is provided in Chapter 18. Additional hammer guidance may be found in The 

Performance of Pile Driving Systems by Rausche et al. (1986), as well as in the 

Deep Foundation Institute Pile Inspector's Guide to Hammers (1995).

Figure 15-24 Pile hammer classification.
356

Table 15-1 Typical Pile Hammer Characteristics and Uses

Hammer Type Single Acting

Air / Steam

Double Acting

Air / Steam

Differential 

Acting 

Air / Steam

Single Acting

Diesel 

(open end)

Double Acting 

Diesel

(closed end)

Rated energy 

(ft-kips) 7 to 1800 1 to 21 15 to 50 9 to 1620 5 to 73

Impact velocity 

(ft/sec) 8 to 16.5 15 to 20 13 to 15 10 to 16.5 8 to 16.5

Blow Rate 

(blows/minute) 35 to 60 95 to 300 98 to 300 40 to 60 80 to 105

Energy 

(per blow)

Ram weight x 

stroke.

(Ram weight + 

effective piston 

head area x 

effective fluid 

pressure) x 

stroke.

(Ram weight + 

effective piston 

head area x 

effective fluid 

pressure) x 

stroke.

Ram weight x 

stroke.

(Ram weight + 

bounce chamber 

pressure) x 

stroke.

Lifting power Air or steam. Air or steam. Air or steam.

Provided by the 

combustion of 

injected diesel 

fluid.

Provided by the 

combustion of 

injected diesel 

fluid.

Maintenance

More complex 

than for drop 

hammer.

More complex 

than for single 

acting.

More complex 

than for single 

acting.

More complex 

than most air 

impact 

hammers.

More complex 

than most air 

impact 

hammers.

Hammer 

suitability for 

types of piles

Versatile for 

any pile, 

particularly 

large concrete 

and steel pipe.

Timber, steel H 

and pipe piles.

Timber, steel H 

and pipe piles.

All types of 

piles.

All types of 

piles.

Major 

advantages

Relatively 

simple and 

moderate cost.

More productive 

than single 

acting. Some 

double acting 

fully enclosed 

and can be 

used 

underwater.

More productive 

than single 

acting. 

Differential air

hammer uses 

less volume of 

air or steam 

than double 

acting and has 

lower impact 

velocity.

Carry their own 

fuel from which 

power is 

internally 

generated. 

Stroke is a 

function of pile 

resistance.

Carry their own 

fuel from which 

power is 

internally 

generated.

Stroke is a 

function of pile 

resistance.

Major 

disadvantages

Need air 

compressor or 

steam plant. 

Heavy 

compared with 

most diesel 

hammers.

Costs more 

than single 

acting. Need air 

compressor or 

steam plant. 

Costs more 

than single 

acting. Need air 

compressor or 

steam plant. 

Heavy 

compared to 

diesel hammer.

Pollution from 

diesel exhaust. 

Higher cost 

hammer. Low 

blows per 

minute at higher 

strokes for 

single acting.

Pollution from 

diesel exhaust. 

Higher cost 

hammer. 

Hammer lift off 

in hard driving 

conditions. 

Remarks

Commonly 

available 

hammer type.

Ram 

accelerates 

downward 

under pressure.

Ram 

accelerates 

downward 

under pressure.

Most commonly 

used hammer 

type. Variable 

stroke.

Limited 

availability and 

use in practice.
357

Table 15-1 Typical Pile Hammer Characteristics and Uses (Continued)

Hammer Type Drop Single Acting

Hydraulic 

Double Acting

Hydraulic Vibratory Resonant

Rated energy

(ft-kips) 7 to 60 25 to 2162 25 to 2581 ---- ---- 

Impact velocity 

(ft/sec) 23 to 33 5 to 18 5 to 23 ---- ---- 

Blow Rate 

(blows/minute) 4 to 8 30 to 50 40 to 90 750 to 2,400 vibrations/minute

up to 10,800 

vibrations/minute

Energy 

(per blow)

Ram weight x 

height of fall.

Ram weight x 

stroke.

(Ram weight + 

effective piston 

head area x 

effective fluid 

pressure) x 

stroke.

---- ---- 

Lifting power

Provided by 

hoisting engine 

or a crane.

Hydraulic Hydraulic Hydraulic or 

electric power. Hydraulic

Maintenance Simple

More complex 

than other 

impact 

hammers.

More complex 

than other 

impact 

hammers.

Highest 

maintenance

cost.

More complex 

than other 

hammer types

Hammer 

suitability for 

types of piles

All types except 

concrete piles

All types of 

piles.

All types of 

piles.

End bearing 

steel H and pipe 

piles. Very 

effective in 

granular soils.

Steel H piles and 

pipe piles.

Major 

advantages

Lowest initial 

cost equipment.

Fully variable 

energy can be 

delivered.

Energy is 

variable over a 

wide range. 

Can be used for 

underwater 

driving.

Can be used for 

pulling or driving. 

Fast operating

pile installation 

tool.

Fast pile 

installation. Very 

low ground 

vibrations. 

Reduced noise.

Major 

disadvantages

Very high 

dynamic forces 

and danger of 

pile damage. 

Lowest pile 

productivity.

High initial cost.

Energy readout 

device 

recommended 

to monitor 

performance.

High initial cost.

Fully enclosed, 

need energy 

readout device 

to monitor 

performance.

High investment 

and 

maintenance. 

Not recommend 

for friction pile 

installations.

Limited 

availability at 

present.

Remarks Generally 

obsolete.

Commonly 

available 

hammer type.

Commonly 

available 

hammer type. 

Variable moment 

hammers can 

help control 

construction 

vibrations.

Newer hammer 

type and may 

require 

additional field 

inspection 

and/or testing.

Note: Vibratory and resonant hammer rated based on power, not energy.
358

15.8.1 Hammer Energy Concepts

Before the advent of computers and the availability of the wave equation to evaluate 

pile driving, driving criteria for a given nominal resistance was evaluated by concepts 

of work or energy. Work is done when the hammer forces the pile into the ground a 

certain distance. The hammer energy was equated with the work required, defined 

as the nominal geotechnical resistance times the final set. This simple idea led 

engineers to calculate energy ratings for pile hammers and resulted in numerous 

dynamic formulas which ranged from very simple to very complex. Dynamic 

formulas have generally been replaced by more reliable methods of resistance 

determination. However, the energy rating legacy for pile hammers remains.

The energy rating of hammers operating by gravity principles only (drop, single 

acting air/steam or hydraulic hammers) was assigned based on their potential 

energy at full stroke (ram weight, w, times stroke, h). Although single acting (open 

end) diesel hammers could also be rated this way, some manufacturers used other 

principles for energy rating. Historically, these hammers have usually been rated by 

the manufacturer's rating, while the actual observed stroke was often ignored in 

using the dynamic formula. In current practice, the stroke is often measured 

electronically from the blow rate, which is an improvement over past practice. In the 

case of all double acting hammers (air/steam, hydraulic, or diesel), the net effect of 

the downward pressure on the ram during the downstroke is to increase the 

equivalent stroke and reduce time required per blow cycle. The equivalent stroke is 

defined as the stroke of the equivalent single acting hammer yielding the same 

impact velocity. The manufacturers generally calculate the potential energy 

equivalent for double acting hammers.

Ideally, the impact velocity, Vi, could be directly computed using basic laws of 

physics from the equivalent maximum stroke as shown in Equation 15-1. 

i = 2ghV Eq. 15-1 

Where:

Vi = impact velocity (ft/s).

g = acceleration due to gravity (32.17 ft/s2

).

h = ram stroke (feet).

The kinetic energy could be computed from Equation 15-2.

2

2

1 KE = mVi Eq. 15-2 
359

Where:

KE = kinetic energy (ft-lbs).

Vi = impact velocity (ft/s).

m = ram mass (slugs/kips). 

If there were no losses, the kinetic energy would equal the potential energy. In 

reality however, energy losses occur due to a variety of factors (friction, residual air 

pressures, preadmission, gas compression in the diesel combustion cylinder, preignition, etc.) which result in the kinetic energy being less than the potential energy. 

It is the inspector's task to identify these losses when possible, and the contractor’s 

task to correct situations where losses are excessive. Some hammers, such as 

modern hydraulic hammers, measure the velocity near impact and hence can 

calculate the actual kinetic energy available.

Further losses occur in the transmission of energy to the pile. The hammer cushion, 

helmet, and pile cushion all have kinetic energy and store some strain energy, while 

the pile head also has inelastic collision losses. Energy is transferred to the pile with 

time and therefore the energy delivered to the pile can be calculated from the work 

done as the integral of the product of force and velocity with time. This is referred to 

as the transferred energy or ENTHRU.

Pile length, stiffness and resistance influence the energy delivered to the pile. The 

actual stroke (or potential energy) of diesel hammers depends on the pile resistance 

and the net transferred energy, which can vary. The stroke of single acting 

air/steam hammers is somewhat dependent upon the pile resistance and rebound 

while the stroke of all double acting hammers is even more dependent on pile 

resistance due to lift-off considerations. In reality, transferred energy increases only 

when both the force and velocity are positive (compression forces; downward 

velocity). As resistance increases and/or for shorter pile lengths, the rebound or 

upward velocity occurs earlier, and the pile then transfers energy back to the driving 

system. In fact, the energy returning to the hammer may occur before all the energy 

has been transferred into the pile.

15.9 DROP HAMMERS

The most rudimentary pile hammer still in use today is the drop hammer as shown in 

Figure 15-25. These hammers consist of a hoisting engine having a friction clutch, a 

hoist line, and a drop weight. The hammer stroke is widely variable and often not 

very precisely controlled. Operation proceeds by engaging the hoist clutch to raise 

the drop weight or ram. The hoist clutch is then disengaged, allowing the drop 
360

weight to fall as the hoist line pays out. Efficiency of the fall is low since the ram is 

attached by a cable to the hoist and must overcome the rotational inertia of the hoist. 

Ideally, the crane operator engages the clutch immediately after impact to prevent 

excessive cable spooling. If the operator prematurely engages the clutch, or it is 

partially engaged during spooling, the fall efficiency and impact energy is reduced.

The hammer operating speed (blows per minute) depends upon the skill of the 

operator and the height of fall being used, but is generally very slow. One of the 

greatest risks in using a drop hammer is overstressing and damaging the pile. Pile 

stresses are generally increased with an increase in the impact velocity (hammer 

stroke) of the striking weight. Therefore, the maximum stroke should be limited to 

those strokes where pile damage is not expected to occur. In general, drop 

hammers are not as efficient as other impact hammers but are inexpensive and 

simple to operate and maintain. Current use of these hammers is generally limited 

to sheet pile installations where pile resistance is not an issue. Drop hammers are 

not recommended for foundation pile installations.

Figure 15-25 Typical drop hammer.
361

15.10 SINGLE ACTING AIR/STEAM HAMMERS

Single acting air/steam hammers are essentially gravity, or drop hammers, for which 

the hoist line has been replaced by a pressurized medium, being either steam or air. 

While originally developed for steam power, the vast majority of these hammers 

today operate on compressed air. To lift the ram weight with motive pressure, a 

simple one cylinder steam engine principle is used. The ram consists of a compact 

block with a so called ram point attached at its base. The ram point strikes against a 

striker plate as illustrated in Figure 15-26. A photograph of a typical single acting 

air/steam hammer is presented in Figure 15-27(a).

During the upstroke cycle, the ram is raised by externally produced air or steam 

pressure acting against a piston housed in the hammer cylinder. The piston in turn is 

connected to the ram by a rod. Once the ram is raised a certain distance, a valve is 

activated and the pressure in the chamber is released. At that time, the ram has 

some remaining upward velocity that depends upon the pile rebound, inlet air 

pressure, and volume of air within the hammer cylinder. Against the action of gravity 

and friction, the ram then "coasts" up to the maximum height (stroke). The 

maximum stroke, and hence hammer potential energy, is therefore not constant and 

depends upon the pressure and volume of air or steam supplied, as well as the 

amount of pile rebound due to soil resistance effects. During the downstroke cycle, 

the ram falls by gravity (less friction) to impact the striker plate and hammer cushion. 

Just before impact, the pressure valve is activated and pressure again enters the 

cylinder. This occurs approximately 2 inches before impact, but depends on having 

the correct hammer cushion thickness. If the hammer cushion height is too low, 

then the pressure is introduced too early, reducing the impact energy of the ram. 

This condition is referred to as "preadmission."

The dynamic forces exerted on a pile by a single acting air/steam hammer are of the 

same short time duration as those exerted by a drop hammer. Because operating 

strokes are generally shorter, the accelerations generated by single acting air/steam 

hammers do not reach the magnitude of drop hammers. Some hammers may be 

equipped with two nominal strokes, one full stroke and another of lesser height. The 

hammer operator can switch between the two to better match the driving conditions 

and limit driving resistance or control tension driving stresses as needed. The 

maximum stroke of single acting air/steam hammers generally ranges from 3 to 5 

feet. The weights of single acting air/steam hammer rams are usually considerably 

higher than drop hammer weights. Single acting air/steam hammers have the 

advantages of moderate cost and relatively simple operation and maintenance. 

They can be used for many pile types, particularly large concrete and steel piles.
362

Figure 15-26 Schematic of a single acting air/steam hammer.
363

Figure 15-27 (a) Single acting air, and (b) differential acting air hammers. 

15.11 DOUBLE ACTING AIR/STEAM HAMMERS

The working principle of a double acting air/steam hammer is illustrated in Figure 

15-28. For a double acting hammer, the ram is raised by pressurized air or steam 

during the upstroke. As the ram nears the maximum up stroke, the lower air valve

opens, allowing the lower cylinder chamber to release the pressurized air. Once the 

ram reaches full stroke, the upper valve changes to admit pressurized air or steam 

to the upper cylinder. Gravity and the upper cylinder pressure accelerate the ram 

through its downward fall. As with the single acting hammer, the stroke is again not 

constant, due to variable lift pressure and volume as well as differing pile rebound. 

During hard driving with high pile rebound, the pressure may need to be reduced to 

prevent lift off, the hammer actually lifting up and away from the pile head. Since the 

maximum stroke is limited and the same pressure is applied during downstroke, a 

reduction in the operating pressure due to lift off may cause the kinetic energy at 

impact to be reduced during these hard driving situations. Just before impact, the 

valve positions are reversed and the cycle repeats.
364

The correct cushion thickness is extremely important for the proper operation of the 

hammer. If the hammer timing is off significantly, it is possible for the hammer to run 

with the ram moving properly, but with little or no impact force delivered to the pile.

The kinetic energy of the ram at impact depends on the ram weight and stroke as 

well as the motive pressure effects. The overall result is that a properly operating 

double acting hammer with its shorter stroke delivers comparable impact energy per 

blow at up to about two times the blow rate of a single acting hammer of the same 

ram weight.

Some double acting air/steam hammers are fully enclosed and can be operated 

underwater. They may be more productive than single acting hammers, but are 

more dependent upon the air pressure. Experience has shown that on average, 

they are slightly less efficient than equivalently rated single acting hammers. Double 

acting hammers generally cost more than single acting hammers and require 

additional maintenance. Similar to single acting air/steam hammers, they require an 

air compressor or a steam plant. However, double acting air/steam hammers 

consume more air and require greater air pressures than equivalent single acting 

hammers. The use of double acting air/steam hammers has diminished and they 

are infrequently encountered in practice.
365

Figure 15-28 Schematic of a double acting air/steam hammer.
366

15.12 DIFFERENTIAL ACTING AIR/STEAM HAMMERS

The differential acting air/steam hammer is another type of double acting hammer 

with relatively short stroke and fast blow rates. A photograph of a differential acting 

air hammer is presented in Figure 15-27(b), while the working principle is illustrated 

in Figure 15-29. Operation is achieved by pressure acting on two different diameter 

pistons connected to the ram. At the start of the cycle, the single valve is positioned 

so that the upper chamber is open to atmospheric pressure only and the lower 

chamber is pressurized with the motive fluid. The pressure between the two pistons 

has a net upward effect due to the differing areas, thus raising the ram. The ram 

has an upward velocity when the valve position changes and applies air pressure 

into the upper chamber, causing the net force to change to the downward direction. 

Thus air pressure along with gravity and friction slows the ram, and after attaining 

the maximum stroke of the cycle, assists gravity during the downstroke to speed the 

ram.

As with the double acting hammers, the kinetic energy at impact may need to be 

reduced during hard driving since the pressure, which assists gravity during 

downstroke, must be reduced to prevent hammer lift-off. As with the other air/steam 

hammers, when the ram attains its maximum kinetic energy just before impact, the 

valve position is reversed and the cycle begins again. Therefore, the hammer 

cushion must be of the proper thickness to prevent preadmission which could cause 

reduced transferred energy. Very high air pressures between 120 and 140 psi at the 

hammer inlet are required for proper operation. However, most air compressors 

only produce pressures of about 120 to 130 psi at the compressor. As with the 

double acting hammer, the efficiency of a differential hammer is somewhat lower 

than the equivalent single acting air/steam hammer. The heavier ram of the 

differential acting hammer is lifted and driven downward with a lower volume of air or 

steam than is used by a double acting hammer. The use of differential acting 

air/steam hammers has also diminished and they too are infrequently encountered in 

practice.
367

Figure 15-29 Schematic of a differential acting air/steam hammer.
368

15.13 SINGLE ACTING (OPEN END) DIESEL HAMMERS

The basic distinction between all diesel hammers and all air/steam hammers is that, 

whereas air/steam hammers are one cylinder engines requiring motive power from 

an external source, diesel hammers carry their own fuel from which they generate 

power internally. Figure 15-30 shows the working principle of a single acting diesel 

hammer. The initial power to lift the ram must be furnished by a hoist line or other 

source to lift the ram upward on a trip block. After the trip mechanism is released, 

the ram guided by the outer hammer cylinder falls under gravity. As the ram falls, 

diesel fuel is injected into the cylinder below the air/exhaust ports. Once the ram 

passes the air/exhaust ports the diesel fuel is compressed and heats the entrapped 

air. As the ram impacts the anvil the fuel explodes, increasing the gas pressure. In 

some hammers the fuel is injected in liquid form as shown in Figure 15-30(b), while 

in other hammers the fuel is atomized and injected later in the cycle and just prior to 

impact. In either case, the combination of ram impact and fuel explosion drives the 

pile downward, and the gas pressure and pile rebound propels the ram upward in 

the cylinder. On the upstroke, the ram passes the air ports and the spent gases are 

exhausted. Since the ram has a velocity at that time, the ram continues upward 

against gravity, and fresh air is pulled into the cylinder. The cycle then repeats until

the fuel input is interrupted.

There is no consensus by the various hammer manufacturers on how a single acting 

diesel hammer should be rated. Many manufacturers use the maximum potential 

energy computed simply from maximum stroke times the ram weight. The actual 

hammer stroke achieved is a function of fuel charge, condition of piston rings 

containing the compressed gases, recoil dampener thickness, driving resistance, 

and pile length and stiffness. Therefore, the hammer stroke cannot be fully 

controlled. A set of conditions will generate a certain stroke which can only be 

adjusted within a certain range by the fuel charge. It may not be possible to achieve 

the manufacturer's maximum rated stroke under all conditions. In normal conditions, 

part of the available potential energy is used to compress the gases as the ram 

proceeds downward after passing the air ports. The gases ignite when they attain a 

certain combination of pressure and temperature. Under continued operation, when 

the hammer's temperature increases due to the burning of the gases, the hammer 

fuel may ignite prematurely. This condition, called "pre-ignition", reduces the 

effectiveness of the hammer, as the pressure increases dramatically before impact, 

causing the ram to do more work compressing the gases and leaving less energy 

available to be transferred into the pile.
369

Figure 15-30 Schematic of a single acting diesel hammer.
370

When driving resistance is very low, the upward ram stroke may be insufficient to 

scavenge (or suction) the air into the cylinder and the hammer may not continue to 

operate. Thus, the ram must be manually lifted repeatedly until resistance 

increases. The stroke can be reduced for most hammers by reducing the amount of 

fuel injected. Some hammers have stepped fuel settings while others have 

continuously variable throttles. Other hammers use pressure to maintain fuel flow by 

connecting a hand operated fuel pump to the hammer, which is operated at the 

ground. By adjusting the fuel pump pressure, hammer strokes may be reduced. 

Using the hammer on reduced fuel can be useful for limiting driving stresses. For 

single acting diesel hammers, the stroke is also a function of pile resistance, which 

also helps in limiting driving stresses. This feature is very useful in controlling tensile 

stresses in concrete piles during easy driving conditions. The actual stroke can and 

should be monitored. The stroke of a single acting diesel hammer can be calculated 

from the following formula:

 Eq. 15-3 

Where:

h = ram stroke (feet).

bpm = blows per minute.

Note: This formula is only applicable for calculating the stroke of single acting 

 diesel hammers and not correct for other hammer types.

Diesel hammers may be expensive and their maintenance more complex. Concerns 

over air pollution from the hammer exhaust have also arisen, causing some areas to 

require a switch to kerosene fuel. However, it should be noted that diesel hammers 

burn far less fuel to operate than the air compressor required for an air/steam 

hammer. To address environmental concerns, some diesel hammers can be 

operated using biodiesel fuel and non-petroleum lubricants. One manufacturer has 

also developed a smokeless diesel hammer. Diesel hammers are also considerably 

lighter than air/steam hammers with similar energy ratings, allowing a larger crane 

operating radius and/or a lighter crane to be used. A photograph of a typical single 

acting diesel hammer is shown in Figure 15-31(a). 
371

15.14 DOUBLE ACTING (CLOSED END) DIESEL HAMMERS

The double acting diesel hammer works very much in principle like the single acting 

diesel hammer. The main change consists of a closed cylinder top. When the ram 

moves upward, air is being compressed at the top of the ram in the so called

"bounce chamber" which causes a shorter stroke and therefore a higher blow rate. 

A photograph of a typical double acting diesel hammer is provided in Figure 

15-31(b).

Figure 15-31 (a) Single acting diesel and, (b) double acting diesel hammers.

The bounce chamber has ports so that atmospheric pressure exists as long as the 

ram top is below these ports, as shown in Figure 15-32. Operationally, as the ram 

passes the bounce chamber port and moves toward the cylinder top, it creates a 

pressure which effectively reduces the stroke and stores energy, which in turn will be 

used on the downstroke. Like the single acting hammer, the actual stroke depends 

on fuel charge, pile length and stiffness, soil resistance, and condition of piston 

rings. As the stroke increases, the chamber pressure also increases until the total 

upward force is in balance with the weight of the cylinder itself. Further compression 

beyond this maximum stroke is not possible, and if the ram still has an upward 
372

Figure 15-32 Schematic of a double acting diesel hammer.
373

velocity, uplift of the hammer will result. This uplift should be avoided as it can lead 

both to an unstable driving condition and to hammer damage. For this reason, the 

fuel amount, and hence maximum combustion chamber pressure, should be 

reduced so that there is only a very slight lift-off, or none at all. Most of these 

hammers have hand held fuel pumps connected by rubber hose to control the fuel 

flow. Hammer strokes, and therefore hammer energy, may be increased or 

decreased by the fuel pump pressure.

To determine the energy provided by the hammer, the peak bounce chamber 

pressure in the hammer is read from a bounce chamber pressure gage. The 

hammer manufacturer should supply a chart which correlates the bounce chamber 

pressure gage reading as a function of hose length with the energy provided by the 

hammer. The use of double acting diesel hammers is limited and they are 

infrequently encountered in practice.

15.15 SINGLE ACTING HYDRAULIC HAMMERS

Single acting hydraulic hammers use an external hydraulic power source to lift the 

ram. They can be perhaps thought of as a modern, although more complicated, 

version of air/steam hammers in that the ram weights and maximum strokes are 

similar in size and the ram is lifted by an external power source. Low headroom 

models exist that can be mounted on an excavator while larger models can be 

mounted on specialty / universal rigs or in conventional leads. During operation, 

hydraulic actuators lift the ram which then retracts quickly at predetermined height. 

This fully releases the ram, which falls under gravity, impacting the striker plate and 

hammer cushion located in the helmet. The hydraulic cylinder then lifts the ram 

again and the cycle is repeated. Single acting hammers are classified as such 

because of dependence upon gravity alone to perform the work. Some hammer 

models include hydraulic accumulators that can provide a relatively small double 

acting component. A schematic of a single acting hydraulic hammer is illustrated in 

Figure 15-33(a), while a photograph of a single acting hydraulic hammer is 

presented in Figure 15-34(a).

Most single acting hydraulic hammers allow the ram stroke, blow rate, and dwell 

time (duration ram remains on top of pile following impact) to be continuously varied 

and controlled using a pendant attached to the hydraulic power pack. Very short 

strokes may be used during easy driving to prevent pile run or to minimize tension 

stresses in concrete piles. Higher strokes are available for hard driving conditions.
374

Figure 15-33 Schematic of hydraulic hammers: 

(a) single acting and (b) double acting.
375

Figure 15-34 (a) Single acting hydraulic, and (b) double acting hydraulic hammers.

On many single acting hydraulic hammers, such as the one illustrated in Figure 15-

33a, the stroke can be easily be visually estimated. Most newer hydraulic hammers 

include a built-in monitoring system which determines the ram velocity just before

impact. The ram velocity can be converted to kinetic energy or equivalent stroke. 

Because of the variability of stroke, this hammer monitor should be required as part

of the hammer system. The monitor results should be observed during pile driving 

with appropriate hammer performance notes recorded on the driving log. Hydraulic 

hammers require a dedicated hydraulic power pack, and can be more complex to 

operate and maintain compared to other hammers.

15.16 DOUBLE ACTING HYDRAULIC HAMMERS

Double acting hydraulic hammers cover the same range of manufacturer’s rated 

energy as the single acting hydraulic hammers but can also be significantly larger. 

Figure 15-33(b) presents a schematic of a double acting hydraulic hammer. Double 

acing hammer consist of a ram attached to a piston rod which are entirely enclosed 

within the hammer housing and an external hydraulic source supplies the power to 

the hammer. Oil flows through a supply valve into the hammer and out through a 
376

return valve to the power pack. During hammer operation, the return valve closes, 

which causes the piston to rise. When the piston reaches a predetermined height, 

the supply valve closes and the return valve opens. Pressure is relieved and the 

ram begins its downward stroke accelerated by hydraulic pressure or compressed 

nitrogen gas. The ram then strikes the anvil, which in turn contacts the pile. Steel 

on steel impact occurs between these two parts which results in a relatively high 

energy transfer.

Similar to single acting hydraulic hammers, the ram stroke can be controlled to adapt 

to the driving conditions. However because of the enclosed housing, the stroke 

cannot be visually estimated. Built-in monitoring systems are therefore essential on 

double acting hammers to document hammer performance. A photograph of a 

double acting hydraulic hammer is included in Figure 15-34(b). A significant

advantage of the fully enclosed double acting hydraulic hammers is that they can 

operate underwater. This allows piles to be driven without using a follower or extra 

length pile. Most double acting hydraulic hammers do not have conventional 

hammer cushions and thus generate steel to steel impacts with high efficiency. 

Some models can also be used to drive piles horizontally.

15.17 VIBRATORY HAMMERS

Vibratory hammers use paired counter-rotating eccentric weights to impart a 

sinusoidal vibrating axial force to the pile (the horizontal components of the paired 

eccentric weights cancel). A schematic of a vibratory hammer is presented in Figure 

15-35 and a photograph is included in Figure 15-36. Several types of vibratory 

hammers exist, variable moment vibratory hammers, high frequency vibratory 

hammers, standard vibratory hammers and low frequency vibratory hammers. 

Variable moment and high frequency hammers typically operate at up to 2300 to 

2400 vibrations per minute or 40 Hz. Standard vibratory hammers operate at up to a 

maximum of 1600 vibrations per minute or 26 Hz and low frequency vibratory 

hammers operate up to about 1200 vibrations per minute or 20 Hz.. The lower the 

operating rate (vibrations per minute) of the vibratory hammer the greater the effect 

on the soil and structure. Hence, hammer start up and shut down often produce the 

most construction vibration concerns. Variable moment vibratory hammers are 

attractive in these situations as the hammer can be brought to full operating speed 

with the eccentric moments in neutral thus avoiding critical frequencies for potential 

vibration damage.
377

Figure 15-35 Schematic of a vibratory hammer.

Figure 15-36 Vibratory hammer installing an H-pile (courtesy ICE).
378

Vibratory hammers are rigidly connected by hydraulic clamps to the pile head and 

may be used for either pile installation or extraction. Vibratory hammers typically do 

not require leads, although templates are often required for sheet pile cells. These 

hammers are not rated by impact energy delivered per blow, but instead are 

classified by energy developed per second and/or by the driving force they deliver to 

the pile. The power source to operate a vibratory hammer is usually a hydraulic 

power pack.

Vibratory hammers are commonly used for driving/extracting sheet piles and can 

also be used for installing non-displacement H-piles and open end pipe piles. 

However, it is often difficult to install closed end pipes and other displacement piles 

due to difficulty in displacing the soil laterally at the toe. Vibratory hammers should 

not be used for precast concrete piles because of possible pile damage due to 

tensile and bending stress considerations. Vibratory hammers are most effective in 

granular soils, particularly if submerged. They also may work in silty or softer clays, 

but most experience suggests they are less effective in stiff to hard clays.

Some wave equation analysis programs can simulate vibratory driving. Dynamic 

measurements have also been made on vibratory hammer installed piles. However, 

a reliable technique for estimating nominal resistance during vibratory hammer 

installation has not yet been developed. If a vibratory hammer is used for pile 

installation, confirmation of the nominal resistance by other means is still necessary.

15.18 RESONANT HAMMERS

Resonant pile hammers use a hydraulic piston-cylinder design to generate a high 

magnitude, high frequency, oscillating force. The amplitude and magnitude of the 

force is controlled through a valve that rapidly switches hydraulic oil to alternating 

sides of the piston that oscillates at up to 180 Hz (10,800 vibrations per minute). 

Piles are advanced into the soil using the high frequency vibration and resonance. A 

proprietary algorithm automatically adjusts and optimizes the frequency to maintain 

resonance. Janes (2009) summarized projects where resonant hammers have been 

used to install steel H-piles and pipe piles.

Advantages of the resonant hammer include fast pile installation, very low ground 

vibrations, and reduced noise levels compared to conventional equipment. Typical 

cranes and hydraulic power packs can be used on resonant pile hammer projects. 

Similar to vibratory installed piles, a disadvantage of resonant hammer use is 

nominal resistance verification. A reliable technique for determining the nominal 
379

resistance is not available. Therefore, a conventional impact hammer and 

associated standard nominal resistance verification methods must be employed after 

resonant hammer pile installation is completed. A photograph of a resonant hammer 

is shown in Figure 15-37. 

Figure 15-37 Resonant hammer installing an H-pile

(courtesy of Resonance Technology). 

15.19 IMPACT HAMMER SIZE SELECTION

It is important that the contractor and the engineer choose the proper hammer for 

efficient use on a given project. An impact hammer which is too small may not be 

able to drive the pile to the required resistance, or may require an excessive number 

of blows. On the other hand, an impact hammer which is too large may damage the 

pile. The use of empirical dynamic pile formulas to select hammer energy and size 

is not recommended as this approach incorrectly assumes these formulas result in 

the desired nominal resistance. Results from these formulas become progressively 

worse as both the complexity of the hammers and the required nominal pile 

resistance increase.
380

A wave equation analysis, which considers the hammer cushion pile soil system, is 

the recommended method to determine the optimum hammer size. In general, a 

hammer having a ram weight of 1 to 2% of the required nominal resistance or 

nominal driving resistance, whichever is greater, often yields a good first estimate of 

the necessary hammer size. Table 15-2 also provides approximate minimum 

hammer energy sizes for preliminary equipment evaluation based on ranges of 

nominal resistances. These are generalizations of equipment size requirements that 

should be modified based on pile type, pile loads, pile lengths, and local soil 

conditions. In some cases, such as short piles to rock, a smaller hammer than 

indicated may be more suitable to control driving stresses. This generalized table 

should not be used in a specification. Guidance on developing a minimum energy 

table for use in a specification is provided in Chapter 14. 

Table 15-2 Preliminary Hammer Energy Requirements

Nominal Resistance

(kips) 

Minimum Manufacturer’s

Rated Energy

(ft-lbs)

180 and under 12,000

181 to 300 21,000

301 to 415 28,800

416 to 540 37,600

540 to 600 42,000

15.20 HAMMER KINETIC ENERGY MONITORING

Several pile driving hammers are available from their manufacturers with kinetic 

energy readout devices. These devices typical monitor hard wired proximity 

switches built into the hammer body. The impact velocity and hammer kinetic 

energy are calculated based on the time it takes the ram to travel the distance 

between the proximity switches. These devices also typically provide the hammer 

blow rate, and for open end diesel hammers, the hammer stroke. Examples of 

hammer manufacturer provided devices are presented in Figure 15-38 for an IHC 

hydraulic hammer and Figure 15-39 and Figure 15-40 for a Berminghammer diesel

hammer.
381

Figure 15-38 IHC hydraulic hammer kinetic energy readout panel. 

Any existing hammer can also be retrofitted for these measurements by attaching 

proximity switches to the hammer body. Attachment procedures vary depending 

upon the hammer model. For a diesel hammer, proximity switches are set into two 

1.2 inch diameter smooth bore drill holes in the cylinder wall above the combustion 

chamber. For air/steam or single acting hydraulic hammers, the proximity switches 

are attached to the hammer body. The proximity switches are connected to a 

transmitter mounted on the hammer that sends the impact velocity and kinetic 

energy to a wireless hand held unit. This hand held unit, called an E-Saximeter, can 

also be used to keep a pile driving log if the inspector presses the enter key with 

each passing pile penetration increment. For single acting diesel hammers, the 

hammer stroke can also be calculated and displayed. A photograph of a diesel 

hammer retrofitted for kinetic energy measurements is presented in Figure 15-41

and the readout device is in Figure 15-42. 

Hammers equipped with kinetic energy readout devices provide improved quality 

control and are particularly attractive on large projects or with piles that require a 

large nominal resistance. These devices can detect changes in hammer 

performance over time that may necessitate adjustment to the pile installation 

criterion.
382

Figure 15-39 Proximity switch attachment for Berminghammer diesel hammer

(courtesy Berminghammer).

Figure 15-40 Proximity switches, readout device and driving log trigger switch for 

Berminghammer diesel hammer (courtesy Berminghammer).
383

Figure 15-41 Proximity switches and transmitter on retrofitted diesel hammer.

Figure 15-42 E-Saximeter wireless kinetic energy readout device.
384

15.21 NOISE SUPPRESION EQUIPMENT

Depending upon the hammer and pile type used, noise from impact pile driving 

operations can range from around 80 to 130 dBa. Local ordinances or specification 

may place limits on noise levels that may influence equipment selection or may 

dictate that pile driving noise shrouds be used in order to meet the specified noise 

limits. Manufacturers of a few diesel and hydraulic hammers can provide optional 

noise suppression devices that may reduce the pile driving generated noise by about 

10 dBa. An example of a noise shroud produced by a hammer manufacturer is 

presented in Figure 15-43. 

Greater reductions in pile driving noise have been obtained by combining noise 

abatement techniques on a project. A 20 to 25 dBa reduction was obtained through 

the combined use of shock absorbing cushion material, a hammer exhaust noise 

shroud, application of damping compound to the steel piles, and use of a noise

shroud around the hammer-pile impact area. Hammer and pile type selection can 

also influence the pile driving generated noise and should be considered in the 

design stage of projects in noise sensitive areas.

Figure 15-43 Noise shroud for IHC hydraulic hammer. 
385

15.22 UNDERWATER NOISE SUPPRESSION EQUIPMENT

Bubble curtains are sometimes required when driving piles through water to reduce 

underwater sound waves, shock waves, and overpressures that impact marine 

mammals and fish. In general, overpressure levels greater than 4.4 psi have been 

found to be detrimental. However, the detrimental overpressure level will vary 

depending upon the species of fish, their size, and their maturity level.

Bubble curtain devices use air bubbles to attenuate the pile driving induced pressure 

wave. Bubble curtains can be categorized as bubble rings or bubble walls. A bubble 

ring is typically used around a single pile and typically consists of a high volume air 

compressor a primary feed line, a primary distribution manifold, medium volume 

secondary feed lines, and secondary distribution manifolds. Bubble walls combine 

the features of bubble rings with a sound damping curtain that encapsulates the air 

bubbles. Bubble walls are typically placed around a complete substructure location 

rather than an individual pile.

For a bubble curtain to be effective, the bubble curtain must completely surround the 

pile driving activity. This can sometimes be difficult to accomplish with a bubble ring 

in areas with tides and currents, or when the foundation design includes batter piles.

Bubble rings are sometimes used in conjunction with containment devices such as a 

large diameter pile sleeve, a turbidity curtain, or a cofferdam in these situations. A 

bubble wall system may be more attractive in areas where a bubble ring system 

requires containment devices. A photograph of a pile driven inside bubble ring used 

in conjunction with a containment device is presented in Figure 15-44. Longmuir 

and Lively (2001) presented a case history where use of a bubble ring reduced

overpressures during pile driving from in excess of 22 psi with no mitigation to less 

than 3 psi. 

The WSDOT and FHWA sponsored a full scale test of new underwater noise 

suppression technology developed by the University of Washington and Marine 

Construction Technologies, PBC. The research study developed a patented solution 

for mandrel bottom driving a steel pipe pile, trademarked a Reinwall pile. In the full 

scale field study where 30 inch pipe piles were driven, Reinwall et al. (2014) 

determined that the air gap between the mandrel and conventional steel pipe 

resulted in a 21 to 23 dBa reduction in underwater noise compared to the 3 to 6 dBa 

reduction achieved with a typical bubble curtain.
386

Figure 15-44 Bubble ring with containment device (courtesy of WSDOT).

15.23 FOLLOWERS

A follower is a structural member interposed between the pile hammer and the pile, 

to transmit hammer blows to the pile head when the pile head is below the reach of 

the hammer. This occurs when the pile head is below the bottom of leads. 

Followers are sometimes used for driving piles below the deck of existing bridges, 

for driving piles underwater, or for driving the pile head below grade. A photograph 

of a follower for driving steel H-piles underwater is presented in Figure 15-45(a).

Maintaining pile alignment, particularly for batter piles, is a problem when a follower 

is used while driving below the bottom of the leads. The use of a follower is 

accompanied by a loss of effective energy delivered to the pile due to compression 

of the follower and losses in the connection. This loss of effective energy delivered 

to the pile affects the blow count to obtain the required nominal resistance. These 

losses can be estimated by an extensive and thorough wave equation analysis, or 

field evaluated by dynamic measurements. In Figure 15-45(b), the hammer-followerpile alignment issues are apparent for the H-pile being driven on a batter.
387

A properly designed follower should have about the same stiffness (per unit length) 

as the equivalent length of pile to be driven. Followers with significantly less 

stiffness should be avoided. Followers often require considerable maintenance. In 

view of the difficulties that can be associated with followers, their use should be 

avoided when possible. For piles to be driven underwater, one alternative is to use 

a hammer suitable for underwater driving.

Figure 15-45 (a) Follower and (b) follower in use drving H-piles. 
388

15.24 JETTING

Jetting is the use of water or air to facilitate pile penetration by displacing the soil. In 

some cases, a high pressure air jet may be used in combination with water. Jets 

may be used to create a pilot hole prior to or simultaneously with pile placement.

Jetting pipes may be located either inside or outside the pile. Jetting is usually most 

effective in loose to medium dense granular soils.

Jetting is not recommended for friction piles because the frictional resistance is 

reduced by jetting. Jetting should also be avoided if the piles are designed to 

provide substantial lateral resistance. For end bearing piles, the final required 

resistance must be obtained by driving (without jetting). Backfilling should be 

required if the jetted hole remains open after the pile installation. A separate pay 

item for jetting should be included in the contract documents when jetting is 

anticipated. Alternatives to jetting include predrilling and spudding.

The use of jetting has been greatly reduced due to environmental restrictions. 

Hence, jetting is rarely used unless containment of the jetted materials can be 

provided. Photographs of a dual jet system mounted on a concrete pile and a 

jet/punch system are presented in Figures 15-46 and Figure 15-47, respectively.

Figure 15-46 Dual jet system mounted on concrete pile (courtesy of Florida DOT).
389

Figure 15-47 Jet/Punch system (courtesy of Florida DOT).

15.25 PREDRILLING

Soil augers or rotary drills may sometimes be used to facilitate driving or limit 

vibrations. Predrilling is sometimes necessary to install a pile through soils with 

obstructions, such as old timbers, boulders, and riprap. Predrilling is also frequently 

used for pile placement through soil embankments and may be helpful to reduce pile 

heave when displacement piles are driven at close spacing.

The predrilled hole diameter depends upon the size and shape of the pile, and soil 

conditions. The hole should be large enough to permit driving but small enough so 

the pile will be supported against lateral movement. The hole size and depth should 

be noted in the contract documents. In most situations, the predrilled hole diameter 

should be 4 inches less than the diagonal of square or steel H-piles, and 1 inch less 

than the diameter of round piling. Where piles must penetrate into or through very 

hard material, it is usually necessary to use a diameter equal to the diagonal width or 

diameter of the piling. When predrilled holes are used in embankments, a hole of up 

to 6 inches larger than the greatest pile cross sectional dimension is sometimes 

used. A separate pay item for predrilling should be included in the contract 

documents when predrilling is anticipated. A photograph of a solid flight auger 

predrilling system is presented in Figure 15-48. 
390

Figure 15-48 Solid flight auger predrilling system.

15.26 SPUDDING

Spudding is the act of opening a hole through dense material by driving or dropping 

a short and strong member and then removing it. The contractor may resort to 

spudding in lieu of jetting or predrilling when the upper soils consist of miscellaneous 

fill and debris. A potential difficulty of spudding is that a spud may not be able to be 

pulled when driven too deep. However, an advantage of spudding is that soil 

cuttings and groundwater are not brought to the ground surface, which could then 

require disposal due to environmental concerns. Two spudding devices are shown 

in Figure 15-49. The spud in Figure 15-49(a) consists of a thick walled pipe section 

with a conical tip formed from steel plates. The spud in Figure 15-49(b) was made 

by adding a wedge shaped tip and plates to a H-pile section.

15.27 EQUIPMENT SUBMITTALS

The Contractor should provide an equipment submittal of all proposed equipment as 

well as the proposed procedures for equipment use to the Engineer prior to pile 
391

Figure 15-49 (a) Pipe pile and (b) H-pile spuds. 

driving operation. In addition, a Pile Installation Plan should be prepared and 

followed throughout the project. This allows the Engineer to evaluate the proposed 

equipment, perform a wave equation analysis with appropriately modeled soil 

conditions, and provide preliminary or final driving criteria. Figure 15-50 presents a 

drive system submittal form covering hammer components and hammer accessories

as well as pile type, pile section, pile splicing and pile toe details. Additional 

information on any jetting, predrilling, or spudding equipment should be submitted in 

conjunction with this hammer submittal form. Reference can be made to Section 

14.2.1 and 14.2.3 of Chapter 14 respectively for specifications regarding the Pile 

Installation Plan and equipment used for driving. 

By submitting technical information related to the pile driving equipment, the 

Engineer can evaluate the suitability and expected performance of the proposed 

system. Without the specific installation system details, the accuracy of any wave 

equation modeling and resulting driving criteria or drivability assessment is reduced. 

An in depth discussion of wave equation analyses is provided in Chapter 12 of this 

manual.
392

Figure 15-50 Drive system submittal form.
393

REFERENCES

Deep Foundations Institute (1981). A Pile Inspector's Guide to Hammers, First 

Edition, Deep Foundations Institute, Englewood Cliffs, NJ.

Deep Foundations Institute (1995). A Pile Inspector's Guide to Hammers. Second 

Edition, Deep Foundations Institute, Englewood Cliffs, NJ.

Janes, M. (2009). Sonic Pile Driving. The History and the Resurrection of VibrationFree Pile Driving. Piledriver Magazine, Q1, Vol. 6, No. 1, pp. 61-67

Longmuir, C. and Lively, T. (2001). Bubble Curtain Systems Help Protect the Marine 

Environment. Pile Driving Contractors Association Summer Newsletter, pp. 

11-16, http://www.piledrivers.org

Passe, Paul D. (1994). Pile Driving Inspector's Manual. State of Florida Department 

of Transportation.

Rausche, F., Likins, G.E., Goble, G.G. and Hussein, M. (1986). The Performance of 

Pile Driving Systems. Inspection Report, FHWA/RD-86/160, U.S. Department 

of Transportation, Federal Highway Administration, Office of Research and 

Development, Washington, D.C., 92 p. 

Reinhall, P.G., Dahl, P.H., and Dardis (2014). Attenuation of Noise from Impact Pile 

Driving Using an Acoustic Shield, Journal of the Acoustical Society of 

America, Vol. 135, No. 4, 2388. 
394
395

CHAPTER 16

PILE ACCESSORIES

Pile accessories are frequently used for pile toe protection or to facilitate pile 

splicing. Accessories are available for driven piles that can make pile installation 

easier and faster. They can also reduce the possibility of pile damage and help 

provide a more dependable and permanent support for any structure. Heavier 

foundation loads, pile installations to sloping rock surfaces or into soils with 

obstructions, and longer pile lengths, are project situations where the use of pile toe 

attachments and splice accessories are often cost effective and sometimes 

necessary for a successful installation. However, pile accessories may add 

significant cost to the project and should not be used unless specifically needed. 

Pile toe attachments and splices for timber, steel, concrete and composite piles are 

discussed in this chapter. During driving and under all applicable loading conditions, 

pile toe attachments and splices should develop the required structural resistance.

16.1 TIMBER PILES

Potential problems associated with driving timber piles include splitting and 

brooming of the pile toe and/or pile head, splitting or bowing of the pile body, and 

breaking of the pile during driving. Protective attachments at the pile toe and at the 

pile head can minimize these problems.

16.1.1 Pile Toe Attachments

Timber pile toe protection devices include steel boots or points. The trend toward 

larger pile hammers and higher nominal resistances may result in greater risk of 

damage for timber piles if obstructions are encountered. Figure 16-1 shows two 

types of commonly use pile toe attachments. Timber pile boots are designed to 

cover the entire pile toe without the need for trimming. Timber pile points involve 

trimming the pile toe to fit the point. Both toe protection devices can be secured 

quickly in the field. The toe protection device is the device over the pile toe and then 

nailing the shoe straps to the pile toe and anchoring the attachment. Figure 16-2

shows a pile boot attached to a timber pile.
396

Figure 16-1 Timber pile toe attachments: (a) Pile boot and (b) Pile point 

(courtesy Skyline Steel).

Figure 16-2 Timber pile with pile boot attached.
397

16.1.2 Attachments at Pile Head

The American Wood Preservers Institute (Collin 2012), recommends banding timber 

piles with heavy metal strapping at the pile head prior to driving to prevent splitting. 

AASHTO (2010) requires collars, bands, or other devices to prevent splitting and 

brooming when the required nominal driving resistance is more than 200 kips on a 

timber pile. A photograph of a banded timber pile head is shown in Figure 16-3. 

Figure 16-3 Banded timber pile head.

16.1.3 Splices

Timber pile splices are undesirable. It is virtually impossible to develop the full 

bending strength of a spliced timber pile. AASHTO (2010) states that timber piles 

should not be spliced unless specified in the contract documents or in writing by the 

engineer. When timber pile lengths prove insufficient for the required nominal 

resistance, longer piles should be ordered, or the foundation should be redesigned 

using additional piles of the furnished length.
398

16.2 STEEL H-PILES

Steel H-piles are often used to penetrate through hard or problematic upper soil 

layers and terminate in a very dense or hard bearing layer such as rock. The 

bearing layer can be at a depth that requires single or multiple H-pile splices. 

Depending on the nature of the surficial materials, pile toe attachments may also be 

needed to prevent or reduce pile toe damage. Pile toe attachments are discussed in 

detail in Section 16.2.1. Splicing methods for H-piles are described in Section 

16.2.2. The importance of proper welding procedures and weld inspection are 

discussed in Section 16.5.

16.2.1 Pile Toe Attachments

Steel H-piles are generally easy to install due to the non-displacement character of 

the pile. Problems can arise when driving H-piles through man-made fills, very 

dense gravel, or deposits containing boulders. If left unprotected under these 

conditions, the pile toe may deform to an unacceptable extent and separation of the 

H-pile flanges and web may occur. Figure 16-4(a) and 16-4(b) show an extracted Hpile that was driven without a pile shoe through a deposit containing boulders. In 

Figure 16-4(a), the pile has just been extracted after dynamic test records indicated 

pile toe damage. The extracted H-pile has a boulder wedged between the flanges. 

In Figure 16-4(b), the H-pile has been laid on the ground and the extent of the toe 

damage is apparent including separation of the flange and web.

Commercially manufactured pile shoes can help minimize or prevent toe damage 

problems. Pile shoes are also desirable for H-piles driven to hard rock to facilitate 

distributing high localized contact stresses over a greater cross sectional area. This 

is particularly helpful when piles are to be driven onto sloping rock surfaces. Pile toe 

reinforcement consisting of steel plates welded to the flanges and web are not 

recommended because the reinforcement provides neither protection nor increased 

strength at the critical area of the flange-to-web connection.

Several manufactured H-pile shoes are available in various shapes and styles as 

shown in Figure 16-5 (a through c). Pile shoes should be fabricated from cast steel 

conforming to ASTM A148/A148M (Grade 90-60). These shoes are attached to the 

H piles with fillet welds along the outside of each flange as illustrated in Figure 16-6. 

Most agencies have a standard detail for shoe attachment. In general, the lower 

portion of the H-pile flanges are beveled and a 5/16 inch fillet weld or greater 

depending on flange thickness used to attach the shoe. 
399

Figure 16-4 Damaged H-pile driven without pile toe protection 

(courtesy Wisconsin DOT).

Figure 16-5 Styles of H-pile driving shoes (courtesy Skyline Steel).
400

Figure 16-6 Pile shoe welded to pile toe using fillet welds along exterior flanges.

Manufacturers recommend different shoe shapes for various applications. It is 

recommended that for a given set of subsurface conditions, pile shoes from different 

manufacturers should be considered as equivalent provided they are manufactured 

from similar strength materials using similar fabrication techniques. Minor variations 

in configuration should be given minimum importance, except in specific subsurface 

conditions where a certain shape would give a definite advantage.

16.2.2 Splices

H-pile splices are routinely made by using a full penetration groove weld along the 

web and both flanges as shown in Figure 16-7. Commercially manufactured H-pile

splicers as shown in Figure 16-8 are also used. Some agencies also utilize a 

welded square or diamond backer plate detail in conjunction with a welded splice. 

H-pile splices must satisfy all the pile structural resistance requirements in axial 

compression, tension, and bending demanded by the foundation loading conditions. 

The appropriate splice details; a full penetration weld or an H-pile splicer, will be 

determined by the foundation loading conditions. Welded splice 
401

Figure 16-7 Full penetration groove weld with backer bars. 

Figure 16-8 H-pile splicer.
402

strength, beyond that required to survive pile driving induced stresses, is necessary 

for piles subjected to bending stresses and cyclic loading. Hence, it is essential that 

H-pile splices be performed by certified welders following approved procedures. 

Further discussion on field welds is discussed in Section 16.5.

For the manufactured H-pile splicer, a notch is cut into the web of the driven pile 

section and the splicer is slipped over that pile. A 5/16 inch fillet weld extending 

vertically from the four corner of the H-pile splicer is then made between the edge of 

the H-pile splicer and the H-pile flanges. The length of each fillet weld is typically on 

the order of 2.5 to 4 inches. The flanges on the add-on H-pile section are chamfered 

at 45 degrees to facilitate flange welding. Typically the add-on section of pile is 

positioned, aligned, and held while a 5/16 inch fillet weld is made between the edge 

of the H-pile splicer plate and the H-pile flanges. The length of each fillet weld is 

again on the order of 2.5 to 4 inches. After completion of the upper welds, either a 

full penetration or partial penetration groove weld (depending on design) is made 

along both flanges. A partially completed splice is shown in Figure 16-9. 

Figure 16-9 Partially completed splice using H-pile splicer.
403

Most H-piles produced today are typically made from ASTM A-572, Grade 50 steel 

or higher. When H-pile splicers are used they should also be made from the same 

high strength steel as the H-pile.

16.3 STEEL PIPE PILES

Steel pipe piles can be designed to be driven either closed end or open end. Pile 

toe attachments are designed for both installation situations. Toe attachments can 

be used to minimize the potential for toe damage in difficult subsurface conditions as 

well as to help limit pile deflection. Steel pipe piles lengths can be easily increased 

in the field using welded or friction splices. Pile toe attachments are discussed in 

detail in Section 16.3.1. Splicing methods for pipe piles are described in Section 

16.3.2. The importance of proper welding procedures and weld inspection are 

discussed in Section 16.5.

16.3.1 Pile Toe Attachments

Pile toe attachments on pipe piles are used to reduce the possibilities of pile toe 

damage and limit pile deflection. Problems during installation of closed end pipe 

piles may arise when driving through materials containing obstructions. In this case, 

piles may deflect and deviate from their design alignment to an unacceptable extent.

When driving open end pipe piles through or into very dense materials, the toe of the 

pile may be deformed.

Closed end pipe piles are most frequently installed with a flat plate welded to the pile 

toe as shown in Figure 16-10. The flat plate thickness typically ranges from 0.75 to 

1.0 inch thick for closed end pipe piles having an outside diameter of 18 inches or 

less. A thicker closure plate may be required by the engineer for larger diameter 

piles. The diameter of the closure plate is typically 5/8 inch larger in diameter than 

the outside diameter of the pipe pile in order to allow a 5/16 inch fillet weld to attach 

the closure plate to the pile toe. The closure plate can also be the same diameter as 

the pile in which case either the pile toe or the closure plate must be beveled so that 

a partial penetration weld can be made. End plates should be made of ASTM

A36/A36M steel or better.

A conical toe attachment can also be used as a pile toe closure device for closed 

end pipe piles. Conical toe attachments can have a rounded shape as shown in 

Figure 16-11(a) or a 60 degree point as shown in Figure 16-11(b). These conical
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Figure 16-10 Closure plate typical of closed end steel pipe pile.

attachments generally cost more than flat closure plate devices. In addition, flat 

closure plates generally develop a higher unit toe resistance. In the same soil 

profile, a closed end pipe pile with a conical toe attachment may drive slightly longer 

than a pipe with a flat closure plate for the same nominal resistance. A conical toe 

attachment is typically welded to the pile toe using a full or partial penetration groove 

weld. Conical toe attachments should be made from cast steel conforming to ASTM 

A148/A148M (Grade 90-60). A photograph of a conical toe attachment welded to a 

pipe pile is presented in Figure 16-12. 
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Figure 16-11 Conical toe attachments for closed end pipe piles: (a) Rounded 

(b) Pointed conical tip (courtesy Skyline Steel). 

 

Figure 16-12 Conical toe attachment welded on a closed end steel pipe pile.

A “rock crusher” driving shoe is a preferred pipe pile closure device where sloping 

bedrock is encountered. This pile toe attachment consists of a thick flat plate with 
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heavy steel plates forming a 45 degree point. A photograph of the rock crusher toe 

protection device is included in Figure 16-13. 

Figure 16-13 Rock crusher driving shoe welded on a closed end steel pipe pile.

When installing open end pipe piles in dense gravel or to rock, the use of cutting 

shoes will help protect the piles and distribute high localized contact stresses over a 

larger pile area. Cutting shoes are made from cast steel conforming to ASTM 

A148/A148M (Grade 90-60). Both inside flange and external flange cutting shoe 

designs are available as shown in Figure 16-14. 

Both cutting shoes have a bearing ring where the open end pipe pile sits on the 

cutting shoe. Cutting shoes with an outside flange can make drilling through the pile 

toe easier, if needed. However, the shoe perimeter is larger in diameter than the 

pile section and can therefore reduce shaft resistance. Both inside and outside 

flange cutting shoes can be welded to pile toe, however, the outside flange is 

sometimes used with only a friction fit. Photographs of inside flange and outside 

flange cutting shoes on open end pipe piles are presented in 16-15 and 16-16

respectively.
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Figure 16-14 Cutting shoes for open end pipe piles: 

(a) Inside flange and (b) Outside flange (courtesy Skyline Steel). 

Figure 16-15 Inside flange cutting shoe welded to open end steel pipe pile.
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Figure 16-16 Outside flange cutting shoe on open end steel pipe pile.

Large diameter open end pipe piles, 36 inches in diameter or greater, are frequently 

used in heavily loaded foundations having a limited number of piles. In these 

situations, use of a thicker wall bottom section over the lower two pile diameters is 

the preferred toe protection mechanism. The thicker wall toe section functions as an 

inside cutting shoe by having the same outside diameter as the design pile section. 

The thicker wall section helps distribute localized stresses and reduce the potential 

for pile toe damage that may otherwise develop due to partial pile toe contact with 

boulders, sloping bedrock, or on batter pile installations.

16.3.2 Splices

Steel pipe pile sections can be spliced by full penetration welding, or by using a 

mechanical drive-fit or friction splicer. Full penetration groove welds are depicted in 

Figure 16-17(a). A schematic of a friction splicer is shown in Figure 16-17(b). When

a friction splicer is used in cases where the full bending strength of the pile is 

needed, fillet welds are required. However, the friction splicer must be fully seated 

into both top and bottom pile sections before performing the fillet weld to avoid 

subsequent cracking of the weld. Therefore, a full penetration weld is preferred in 

cases where the full bending strength is required.

For the full penetration weld, a backing ring, shown in Figure 16-18, is designed to 

sit between both pile sections and aides in pile alignment. Pins extending from the 
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Figure 16-17 Splices for pipe piles.

Figure 16-18 Backing rings for pipe pile splicing

(a) with pins (courtesy Skyline Steel) and (b) without.
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backing ring provide a gap to facilitate welding the root pass. A photograph of the 

backup ring in place on a lower pile section is shown in Figure 16-19. Following 

completion of the root pass, additional passes are made around the splice to 

complete the full penetration weld and connect the pipe pile sections. Figure 16-20

shows a completed splice.

Figure 16-19 Backing ring tack welded to bottom pile.
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Figure 16-20 Welded pipe pile sections.

In applicable situations, pipe piles can also be spliced with manufactured drive-fit or 

friction splicers similar to the one shown in Figure 16-21. These splicers are 

typically cast from ASTM A27 (Grade 65-35) or ASTM A148/A148M (Grade 90-60) 

steel and are designed with a taper for a frictional connection to eliminate the need 

for welding. A 3/8 inch bearing is located at the midpoint of the friction splicer for 

load transfer from the top pipe section to lower pile section. Little advanced 

preparation is required for this splice; however the adjoining pile sections must be 

square. If the initial pile section has some pile top damage following driving, the 

damaged section must be cut off and made square for a suitable friction connection 

with the drive-fit splicer. Unless a friction splicer is fillet welded to the top and 

bottom pile sections, the full pile strength in bending will not be provided. The 

suitability and location of friction splicers on piles subject to uplift loads should also 

be determined by the design engineer. If friction splicers are used on spiral welded 

pipe piles, fillet welds are also often necessary to provide a barrier to prevent ground 

water from leaking into the pile. A completed splice using the drive-fit friction splicer 

is shown in Figure 16-22. 
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Figure 16-21 Drive-fit mechanical friction splicer for pipe pile

(courtesy Skyline Steel). 

Figure 16-22 Drive-fit mechanical friction splicer on pipe pile.
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16.3.3 Constrictor Plates

As discussed in Section 6.4.2 and Section 7.10 7, large diameter open ended pipe 

(LDOEP) piles can “core” or remain unplugged during driving. To facilitate plugging 

at a targeted depth in the soil stratigraphy, a constrictor plate is sometimes designed 

and welded inside the pile. Figure 16-23 shows the top surface of a constrictor plate 

with stiffeners welded inside a 60 inch O.D. pipe pile. A different constrictor plate 

detail is presented in Figure 6-14 of Section 6.4.2. Constrictor plates are often used 

inside large diameter open end pipe piles to force plugged behavior and a larger 

nominal resistance. Research on LDOEP pile design including constrictor plates is 

ongoing.

Figure 16-23 Constrictor plate installed inside pipe pile.
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16.4 PRECAST CONCRETE PILES

16.4.1 Pile Toe Attachment

The toe of precast concrete piles may be damaged in compression under hard 

driving. In hard driving conditions and for toe bearing piles on rock, special cast 

steel toe attachments can be added to the pile during casting. However, toe 

attachments are not routinely used for concrete piles.

When necessary, a flat cast steel shoe as shown in Figure 16-24(a), or an "Oslo 

Point" or “Rock Injector Point” as shown in Figure 16-24(b), can be cast into a 

concrete pile for toe protection. The characteristics of Oslo or Rock Injector points

are such that the pile can be “chiseled into” most rock types to ensure proper 

seating. These points generally have a 3.5 inch diameter hardened steel point 

housed in the steel casting attached to the concrete pile. All toe attachments for 

precast concrete piles must be attached during casting of the piles and not in the 

field. A photograph of an Oslo point with and without the hardened steel point in the 

housing is presented in Figure 16-25. 

Figure 16-24 Pile toe attachments on concrete pile: 

(a) Flat shoe (courtesy Titus Steel) and (b) Oslo point.
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Figure 16-25 Oslo point for concrete pile. 

Another toe attachment sometimes used with a prestressed concrete pile is an Hpile or pipe pile section cast into the pile toe. These devices are not typically used 

for toe protection but more to facilitate overall pile penetration depths in very dense 

or hard materials where it may be difficult to achieve the required pile penetration 

depth with a displacement pile. Depending on the characteristics of the materials to 

be penetrated, the embedded steel sections may or may not be equipped with its’ 

own toe protection attachment. A photograph of a square prestressed concrete pile 

section with an embedded H-pile section at the pile toe is presented in Figure 16-26. 

In this situation the steel section must be appropriately sized for the concrete 

section, foundation loading conditions, and subsurface conditions to prevent 

overstressing.

Figure 16-26 Steel H-pile cast into toe of prestressed concrete pile.
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16.4.2 Splices

Virtually all concrete piles driven in the United States are prestressed to minimize 

potential problems associated with handling and tension stresses during driving. 

However, the ends of prestressed concrete piles are not effectively prestressed due 

to development length, and thus special precautions must be taken when splicing 

prestressed concrete piles. Whenever possible, concrete piles should be ordered 

with sufficient length to avoid splicing. However, if splicing is required, the splices 

available can be divided into four types: Dowel, Welded, Mechanical, and Sleeve. 

Illustrations of these splice types are provided in Figure 16-27. The wedge and 

pinned connectors can be classified as mechanical splices while the connector ring 

is a sleeve splice.

As part of a FDOT research effort on development of a new post tensioned concrete 

pile splice, Mullins and Sen (2015) summarized the available splicing systems for 

prestressed concrete piles including the capacity, failure type, durability, installation, 

and production aspects. The results of this summary are presented in Table 16-1. 

Table 16-1 Summary of Precast Concrete Pile Splices 

(after Mullins and Sen 2015)

Name of Splice Type Capacity Failure 

Type Durability Installation U.S. Produced

Epoxy Dowel Dowel Poor Ductile Good Moderate / 

Poor Yes

Kie-Lock Mechanical Good Ductile Moderate Good No

ICP Welded Good Brittle Moderate Moderate No

NU Chuck Bolted Good Brittle Moderate Untested / 

Moderate Yes

UF Tube Grouted 

Tube Good Ductile Good Poor Yes

GYA Mechanical Good Ductile Moderate Good No

Macalloy Post 

Tensioned Good Ductile Good Moderate / Poor No

In the above table, capacity was defined as how well the splice met the flexural and 

tensile strength requirements for driving forces and pile loads. The failure type 

identified how the splice fails. Ductile failure is preferable to brittle failure. Durability 

noted the splice resistance to reinforcement and strand corrosion. Installation 

identified the time and labor requirements to complete the splice and U.S. produced 
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identifies whether the splice is made in the U.S. Foreign made splices require 

special exceptions on some Federal funded projects. Some splices are only 

applicable for a specific pile type such as the UF Tube which is designed for voided 

concrete piles.

Figure 16-27 Commonly used prestressed concrete pile splices (after PCI 1993).
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A schematic of an epoxy dowel splice for prestressed concrete pile is presented in 

Figure 16-28 while photographs of the splice in progress are presented in Figure 

16-29(a) and 16-29(b). The bottom pile section to be spliced requires holes to 

receive the dowels. These holes may be cast into the pile when splicing is planned, 

or drilled in the field when splicing is needed, but was unexpected. The bottom 

section is driven with no special consideration and the top section is cast with the 

dowel bars in the end of the pile. When spliced together in the field, the top section 

with the protruding dowels is guided and set in position and a thin sheet metal form 

is placed around the splice. Epoxy is then poured, filling the holes of the bottom 

section and the small space between the piles. Pile driving can typically resume on 

epoxy dowel splices one to two days after epoxy placement. Epoxy dowel splices 

can be time consuming but are comparatively inexpensive. These splices have 

been reliable if dowel bars are of sufficient length and strength, and if proper 

application of the epoxy is provided. The number, length, and location of the dowel 

holes, as well as the dowel bar size, must be designed. As noted by Mullins and 

Sen (2015), the splice has limitations meeting all the flexural and tensile strength 

requirements for driving forces and foundation loads.

Figure 16-28 Cement/Epoxy-dowel splice (after Bruce and Herbert 1974).
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Figure 16-29 Epoxy-dowel splice: (a) pile with core holes and (b) splice in progress. 

Welded splices require having steel fittings cast into the end sections of the concrete 

pile when manufactured. Sections are welded around the entire perimeter. Figure 

16-30(a) shows the beveled steel fitting at the head of the driven pile section and 

Figure 16-30(b) shows the completed ICP pile welded splice.

Figure 16-30 Welded ICP splice: (a) beveled steel fitting and (b) welded splice. 

Most mechanical splices for prestressed concrete piles, such as the Kie-Lock, 

Herkules, and ABB, among others, are made of steel castings and are available for 

square, octagonal, hexagonal, and/or round sectional shapes. The steel castings 

are installed in the formwork of a prestressed concrete pile prior to concrete 

placement. The Kie-Lock (Figure 16-31) and Hercules splices require mating both 

male and female castings, while most other mechanical splices are gender neutral. 

All mechanical splices are then locked by inserting bars, wedges, pins, keys, or 

other mechanical connections after aligning the sections. Although mechanical 

splices can be expensive, they do save considerable time and they have been 

designed to properly account for all loading conditions, including tension. 
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Figure 16-31 Kie-Lock splice.

A new concrete pile splice is under development for the Florida DOT in research by 

Mullins and Sen (2015). The splice is shown in a horizontal position in Figure 16-32. 

The splice is post-tensioned after joining the two concrete pile sections with the full 

prestress level restored across the splice and adjacent pile sections. For this splice, 

anchorages for the strands are cast into the lower section and open ducts are cast 

into the upper pile section. Steel strands are then locked into the lower anchor 

blocks and inserted though the upper pile section. After the two pile sections are 

mated, the steel strands are post-tensioned and the remaining duct voids grouted.

Figure 16-32 Post-tensioned splice (a) splice joint and (b) post tensioning strands in 

upper section (courtesy University of South Florida).

Sleeve type concrete splices are illustrated in Figure 16-33. These concrete pile 

splices can be rapidly applied and are very effective in reducing tension driving 

stresses. However, they cannot be used where static uplift loading will be required. 
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A Hawaiian can sleeve splice is shown in Figure 16-33(a). The sleeve must have 

sufficient length and strength if lateral or bending loads are anticipated. A Bruns 

connector ring splice is shown in Figure 16-33(b). The Bruns splice requires steel 

castings at the ends of the concrete piles. The shorter connector ring design has 

limited tensile and flexural strength and is therefore not recommended for designs 

with those loading considerations.

Figure 16-33 Sleeve splices: (a) Hawaiian can and (b) Bruns splice.

If a specific splice is specified based on previous experience, then an option for 

substituting some other concrete splice should not be allowed unless the substitute 

splicer is field tested. The alternative splice should be required to have equivalent 

compressive, tensile and flexural strength to the originally specified splice. The 

substitute splicer can be tested by driving a number of spliced test piles and 

observing the performance.
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16.5 WELDED SPLICES AND TOE ATTACHMENTS

Field welding is required on many pile splices and pile toe attachments. Weld 

quality is paramount in maintaining the full pile material strength. Due to the 

importance of field welding quality, the Michigan Department of Transportation 

developed a detailed document “Field Manual for Pile Welding” (2012). The New 

York State Department of Transportation also devoted twelve pages in their “Pile 

Driving Inspection Manual” (2012) to pile splicing and toe attachments. The 

NYSDOT document includes minimum splice times for weld preparation and welding 

as a function of pile section and splice type.

If not performed properly, welds at splice locations can crack or break during driving. 

Examples of welded pile splices that failed during driving are presented in Figures 

16-34 and 16-35. The pipe pile weld in Figure 16-34 cracked as a result of bending 

stresses occurring during hammer-pile alignment. The weld shown in Figure 16-35

broke a few hammer blows following splicing. The H-pile splice had been completed 

in low temperatures without properly preheating the pile metal. The main sources of 

welded splice problems, when they occur, can be traced to poor pile surface 

preparation, insufficient bevel or root opening, or general non-compliance with 

approved welding procedures or splice details.

Figure 16-34 Cracked weld on pipe pile splice.
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Figure 16-35 Broken weld on H-pile splice. 

16.5.1 Welding Surface Preparation

Prior to commencing welding, the pile joint to be welded must be properly cleaned. 

If the ambient temperature is below 32 degrees Fahrenheit the weld area must also 

be preheated. Figure 16-36 illustrates a steel pipe pile head which has been 

properly cut and beveled at the required angle to accommodate a welded splice. 

Without the appropriate bevel, weld penetration will be reduced, and it will not be 

possible to complete a full penetration groove weld. In Figure 16-37, a grinder is 

being used to prepare the pile head surface for welding. Typically, the add on H-pile

section will have a 45 degree bevel with a 1/8 or 1/4 inch root opening to complete a 

full penetration weld. In Figure 16-38, two H-pile sections have been positioned for 

splicing. The specified splice detail required a full penetration weld. However, a full 

penetration weld is clearly impossible given the lack of a root opening and bevel 

shown in the figure.
424

Figure 16-36 Beveled edges on pipe pile in preparation for weld.

Figure 16-37 Grinding edges of web and flange in preparation for weld.
425

Figure 16-38 H-pile splice preparation without adequate root opening or bevel.

Proper pile section alignment is required prior to splicing. Figure 16-39 illustrates 

poor section alignment, a poor quality groove weld on the flange, and poor quality 

and insufficient fillet weld lengths on a splice with an H-pile splicer. When H-pile

splicers are used, section alignment and root opening should be checked prior to 

welding. Figure 16-40 shows an extracted H-pile from the same project where

dynamic testing indicated a cracked weld during driving. The H-pile was extracted

after driving; however only the upper pile section returned. The pile separated at the 

poorly welded splice with the H-pile splicer and lower pile section lost below ground.

Figure 16-39 Misalignment between pile sections and poor quality weld.
426

Figure 16-40 Extracted H-pile upper section. Weld cracked during driving and 

extracted section retrieved without H-pile splicer or lower pile section.

16.5.2 Temperature Requirements During Welding and Splicing

In colder climates where the ambient air temperature is below 32 degrees 

Fahrenheit, the steel in the vicinity of the splice should be preheated to a minimum 

temperature of 70 degrees Fahrenheit within 3 inches of the weld immediately prior 

to welding and maintained at a minimum of 50 degrees Fahrenheit until the welding 

is complete. No welding should be performed when the ambient air temperature is 

below 0 degrees Fahrenheit. Figure 16-41 illustrates an H-pile being preheated at 

the splice location prior to welding.
427

Figure 16-41 Heating pile splice location prior to welding.

16.5.3 Welding Pile Toe Accessories

Pile toe attachments should be welded to the toes of H-piles and pipe piles following 

the approved welding procedure by approved welders. In Figure 16-42, small spot

welds; one on the web and one at the corner of each flange have been used to 

attach the driving shoe to the pile toe. The limited welding can facilitate the loss of 

the driving shoe when difficult driving conditions are encountered. A partial 

penetration groove weld along the full length of the H-pile flange is shown attaching 

the driving shoe to the H-pile in Figure 16-43. This detail greatly reduces the 

possibility of driving shoe loss or the shoe not performing as intended when cobbles 

or bedrock is encountered.
428

Figure 16-42 Unsatisfactory spot welds used for shoe attachment.

Figure 16-43 Full groove weld between pile and toe attachment.
429

16.5.4 Welded Splice Checklist

The following checklist is from the Michigan Department of Transportation Field 

Manual for Pile Welding (2012). It highlights several key items to be addressed for 

successful field welded splices.

16.5.4.1Preparation

• Welders should be certified by an authorized agency to perform the type of 

work, and produce the type of weld to be used.

• A Weld Procedure Specification (WPS) should be developed and approved 

prior to work commencement.

• Joint preparation, pile alignment, root openings, bevels and fit up should be 

completed according to contract plans.

• Weld surfaces should be ground and cleaned. Coatings, oil, grease, rust, dirt, 

moisture and other contaminants within the weld zone should be removed.

• When ambient air temperature is below 32 degrees Fahrenheit, the pile splice 

area should be preheated, and maintained at 70 degrees. Welding should 

not be performed when ambient temperature is below 0 degrees Fahrenheit. 

Heating and housing may be approved by an engineer.

• The contractor should have an approved quality control plan, and perform 

welding with respect to the established quality control plan.

16.5.4.2During Welding 

• Electrodes should be properly stored. This includes:

oElectrodes should be stored in a hermetically sealed container or hot box

with a minimum temperature of 250 degrees Fahrenheit.

oOne exposed to the atmosphere, electrodes should be either used within 

two-hours, or redried for two hours at a minimum of 500 degrees.

oElectrodes that become wet should be discarded and not used.

oElectrodes should be discarded if they are dropped on the ground, 

exposed to rain or not properly stored.

• Tack welds used during fit up should not be part of the final weld.

• The welder(s) should follow the approved WPS, contractor quality control plan 

and should inspect his/her own work.

• The welder(s) should clean between passes, removing all slag and repairing 

discontinuities between passes.

• The welder(s) should back gouge to sound metal for full penetration groove 

welds before welding the backside (if applicable). 
430

16.5.4.3Final Weld Inspection

• All slag, spatter, and debris should be removed from weld surface.

• The weld size, length and profile should meet project requirements.

• Arc strikes should be ground smooth.

• A final inspection for cracks or other discontinuities, porosity, undercut, 

underfill, overlap, lack of penetration, or lack of fusion should be performed 

before accepting the weld.

• Repairs or replacement of weld should be performed before accepting the 

weld.
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REFERENCES

American Association of State Highway and Transportation Officials (AASHTO). 

(2014). AASHTO LRFD Bridge Design Specifications, US Customary Units, 

Seventh Edition, with 2015 Interim Revisions. American Association of State 

Highway and Transportation Officials, Washington, D.C., 1960 p.

ASTM A148-14. (2014). Standard Specification for Steel Castings, High Strength, for 

Structural Applications. Book of ASTM Standards, Vol. 1.02, ASTM 

International, West Conshohocken, PA, 4 p.

ASTM A27-13. (2014). Standard Specification for Steel Castings, Carbon, for 

General Application. Book of ASTM Standards, Vol. 1.02, ASTM International, 

West Conshohocken, PA, 4 p.

ASTM A572-15. (2015). Standard Specification for High-Strength Low-Alloy 

Columbium-Vanadium Structural Steel. Book of ASTM Standards, Vol. 1.04, 

ASTM International, West Conshohocken, PA, 4 p.

Collin, J.G. (2002), Timber Pile Design and Construction Manual. American Wood 

Preservers Institute (AWPI), 122 p.

Michigan Department of Transportation (MDOT). (2012). Field Manual for Pile 

Welding. Division of Bridge Field Services, First Edition, MI, 35 p.

Mullins, G., and Sen, R. (2015). Investigation and Development of an Effective, 

Economical and Efficient Concrete Pile Splice. Florida Department of 

Transportation, Tallahassee, FL, 270 p.

New York Department of Transportation (NYDOT). (2012). Pile Driving Inspection 

Manual, GEM-26, pp. 13-25.

Precast/Prestressed Concrete Institute (PCI). (1993). Precast/Prestressed Concrete 

Institute Journal. Vol. 38, No. 2, pp.14-41.

Washington State Department of Transportation (WSDOT). (2012). Standard 

Specifications for Road, Bridge and Municipal Construction. WS, 974 p.
432
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CHAPTER 17

DRIVING CRITERIA

Pile foundations must be installed to meet limit state requirements for compressive, 

tensile, and lateral resistances as well as to satisfy serviceability requirements. 

Driven pile foundations are therefore used to satisfy nominal resistance 

requirements, pile penetration requirements, or both. As the pile is driven into the 

soil or rock, observations of the pile penetration resistance or blow count provide an 

indication of the nominal geotechnical resistance. Several methods have been 

developed that correlate the observed blow count and the associated hammer 

operational performance with the nominal geotechnical resistance achieved.

Driven pile foundations are generally installed using a driving criterion that equates 

the blow count and hammer stroke to the nominal geotechnical resistance. A 

minimum pile penetration depth may often be included in the driving criterion in 

cases when scour, foundation settlement, uplift or lateral loading demands impact 

design performance. The correlation of the observed blow count and hammer stroke 

to the nominal geotechnical resistance can be established from one or a 

combination of two of the following methods: static load test (Chapter 9), dynamic 

testing with signal matching (Chapter 10), rapid load testing (Chapter 11), wave 

equation analysis (Chapter 12) or dynamic formulas (Chapter 13).

17.1 DEVELOPMENT OF THE PILE DRIVING CRITERIA

The foundation designer should specify the method to be used for determination of 

the driving criteria. Furthermore, construction personnel should clearly understand 

the method being used and its proper implementation on the project.

Brown and Thompson (2011) summarized current state transportation agency 

practices on driving criteria in the NCHRP Synthesis 418, “Developing Production 

Pile Driving Criteria from Test Pile Data.” This report surveyed 42 transportation 

agencies covering a wide range of pile types and sections, hammer types and sizes, 

project sizes, as well as soil and rock conditions. The agencies reported their 

predominant method of establishing driving criteria were, from most frequent to least 

frequent:
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• Drive piles to a blow count determined from another pile that had previously 

been subjected to dynamic testing and analysis.

• Drive piles to a blow count determined from a dynamic formula.

• Drive piles to a blow count determined from a wave equation analysis.

• Drive piles to practical refusal.

• Drive piles to a specified tip elevation.

• Drive piles to a blow count determined from another pile that had previously 

been statically load tested.

The resistance verification methods mentioned above each have their own 

resistance factor in AASHTO (2014). Hence, a change in the resistance verification 

method during construction will necessitate driving production piles to either a higher 

or lower nominal resistance than shown on the original contract documents. For 

example, a change from dynamic testing with signal matching (φdyn = 0.65) to the 

FHWA modified Gates formula (φdyn = 0.40) will result in a 63% increase in the 

required nominal resistance. This can create problems if the pile driving system was 

not originally sized for the higher nominal resistance. Conversely, a switch from the 

FHWA modified Gates formula (φdyn = 0.40) to either a static load test (φdyn = 0.75) 

on one pile, or to 100% dynamic testing with signal matching will result in an 87% 

reduction in the required nominal resistance. In this scenario, reduced nominal 

resistance piles may allow the use of a smaller pile hammer, or may solve a 

problematic pile installation condition such as avoiding a damage causing boulder 

layer since piles with a reduced nominal resistance may terminate at a higher, and 

therefore, safer installation elevation.

When developing driving criteria, consideration should also be given to the effect of 

time dependent changes in the nominal geotechnical resistance. In conditions 

exhibiting soil setup, pile penetration resistances at the end of driving less than that 

required for the nominal resistance may be acceptable when setup is confirmed by 

later restrikes. Conversely in geomaterials that exhibit relaxation, pile penetration 

resistances at the end of driving should be greater than those needed for the 

required nominal resistance to account for the future loss of nominal resistance.

Restrike tests are typically performed in soils with time dependent soil strength 

changes to confirm the expected change in nominal resistance. Restrike tests are 

also performed on projects in the event unexpected changes in nominal resistance 

occur with time. In cases where time dependent soil strength changes are 

anticipated, static or dynamic restrike tests should be delayed by an appropriate

waiting period until the anticipated soil strength changes have occurred. Additional 
435

information on time dependent soil resistance effects is presented in Section 7.2.4 

and approximate waiting periods for various soil types is noted in Section 14.5.2.1. 

When time dependent soil strength changes occur, care has to be taken to assess 

the driving criterion for the soil condition at the end of driving. For example, if soil 

setup results in a 50% increase of nominal resistance from the end of driving to the 

time of an acceptable setup period, then piles may be driven to a blow count that 

produces the required nominal resistance divided by 1.5. Obviously, this reduced 

driving criterion may result in substantial savings compared to driving the piles to the 

full nominal resistance. Restrike tests should be performed on a representative 

percentage of production piles to substantiate that the anticipated soil setup occurs 

over the site. A contingency plan should also be in place in the event restrike results 

indicate less than required nominal resistance, such as performing a second longer 

term restrike or driving piles to a greater penetration depth.

17.2 PRACTICAL AND ABSOLUTE REFUSAL

As noted above, agencies sometimes use practical refusal as a driving criteria. 

Definitions for practical and absolute refusal are based on an approved hammer 

system operating properly at its maximum fuel or stroke setting unless hammer 

approval was established based on hammer operation at a reduced fuel or stroke 

setting. If refusal driving conditions develop in combination with suspect hammer 

performance, further evaluation of hammer energy transfer and the source of the 

refusal driving conditions are appropriate.

Practical refusal is defined as a pile penetration resistance (blow count) of 10 blows 

per inch for a maximum of 3 consecutive inches of pile penetration. Practical refusal 

is often used as a criterion for piles driven to a consistent and hard bearing layer. 

Blow counts greater than 10 blows per inch should be used with care for concrete 

piles and should be avoided for timber piles. Absolute refusal is defined as 20 blows 

for one inch or less of pile penetration. Driving should terminate immediately once 

either criteria are achieved with a properly sized and properly working hammer.

Practical and absolute refusal criteria should be used to avoid driving for an 

extended duration at excessively high and unreasonable blow count requirements. 

When seating a pile on hard rock, an absolute refusal criterion of 5 blows per 1⁄4 inch 

or 10 blows per 1⁄2 inch may be preferred to 20 blows per inch to reduce the risk of 

pile toe or driving equipment damage.

When the required pile penetration depths cannot be achieved by driving without 

exceeding practical or absolute refusal criteria with the approved hammer, use of 
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other pile penetration aids should be evaluated. Predrilling, jetting, and spudding 

equipment are discussed in Chapter 15.

17.3 PRACTICAL ISSUES AND CONSIDERATIONS

Prior to the onset of test pile or production pile driving, numerous practical issues 

and considerations should be clearly understood by the parties responsible for 

developing the driving criteria, approving a submitted criteria, or responsible for 

implementing the criteria in the field. Some of the more common issues 

encountered include:

• The lack of, or an incomplete or conflicting definition of, pile driving acceptance 

criteria within contract documents.

Specifications should clearly identify the pile acceptance requirements 

(nominal resistance requirements in compression or uplift; required minimum 

pile penetration depth, if any, to meet lateral resistance or serviceability 

requirements, to avoid premature pile acceptance in competent layers 

overlying compressible ones, or due to high driving resistance caused by 

obstructions or poor hammer performance, etc.) as well as the resistance 

verification method. 

• Clear identification of who determines when driving is terminated on test or 

production piles.

On test pile projects, numerous parties may be present on the job site each 

with different responsibilities. A pile “inspector” will likely be present for the 

agency or for the DB team to record the pile penetration resistance or blow 

count as well as the associated hammer performance during test pile 

installation. The agency’s project engineer or its design consultant may be 

present to observe test pile driving, another agency or contractor retained 

engineer may be present performing a dynamic pile test during driving, and 

the contractor’s foreman may be documenting pile installation. Who and what 

(resistance verification method) determines when test pile driving is 

terminated should be clearly established in the contract documents and 

revisited prior to test pile installation. A detailed, manually recorded, pile 

driving log should always be maintained as part of a test pile installation.
437

During production driving, the pile inspector often serves as the agencies lone 

representative during driving and determines when the driving criteria 

established from the test pile program or by other means is achieved.

• Definition, or lack thereof, of time considerations in pile acceptance criterion.

Pile acceptance criteria are based on achieving a nominal resistance at a 

specific time. Therefore, the pile acceptance criterion must clearly state when 

the criterion is to be applied such as at the end of initial driving or at the 

beginning of restrike. If based on restrike conditions, the time window when 

restrike tests are to be performed must be incorporated into the criterion. 

Oftentimes a time dependent nominal resistance correlation is established 

that requires restrike tests within a set time period. Restrike tests performed 

too early or too late may invalidate this correlation. Restrike tests on piles 

having limited drivability in soil setup environments may also need to be 

performed within a set time window to avoid refusal condition during restrike.

• Definition, or lack thereof, of driving criteria applicability and refinement as 

necessary to various substructure locations, subsurface conditions, or nominal 

resistance demands.

Contract documents should identify test pile locations, the substructure units 

covered by those test pile locations, and that the resultant driving criterion is 

applicable to those substructure units. In some cases developing an 

appropriate driving criterion may necessitate the extrapolation of the test pile 

results to higher or lower nominal resistances, shorter or longer pile lengths, 

different batter angles, or slightly variable soil stratigraphy within the identified 

substructure locations by using refined and then modified wave equation 

analysis. 

• The purpose of specifying and achieving a minimum pile penetration depth or 

required pile toe elevation.

Minimum pile penetration requirements may be specified due to the depth of 

scour, the compressibility of soil layers, the presence of liquefaction 

susceptible layers, satisfying lateral or uplift loading demands, and other 

factors besides the nominal geotechnical resistance in axial compression. A 

minimum penetration depth should generally not be specified only to achieve 

an axial compression resistance.
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• The difference between estimated and minimum pile toe elevations and when a 

minimum pile toe elevation should be specified.

The estimated pile toe elevation defines the estimated pile length needed to 

achieve the nominal geotechnical resistance in axial compression. It is used 

to establish contract length estimates for bidding purposes. A minimum pile 

toe elevation is used to satisfy performance requirements.

• The difference between the required nominal geotechnical resistance based on 

design requirements and the resistance required to be overcome during initial 

driving (greater or lesser) based on unsuitable layers or time dependent 

changes in soil resistance.

The required nominal driving resistance, Rndr, is the soil resistance that must 

be overcome at the time of pile driving in order to provide the nominal 

geotechnical resistance, Rn, needed to satisfy loading and performance 

requirements. The nominal driving resistance can be significantly greater 

than the nominal resistance in cases where scourable, liquefiable, highly 

compressible, or otherwise unsuitable soil layers must be penetrated and/or 

where toe resistance may decrease after initial driving due to relaxation. The 

nominal driving resistance can also be less than the nominal resistance in 

cases where these conditions are not present and/or the suitable support 

layers exhibit soil setup.

• Unreasonable definitions of practical and absolute refusal criteria in terms of 

either an excessive driving resistance (blow count) or a hard to achieve 

minimum pile penetration length.

Definitions of practical and absolute refusal with regard to blow count aspects 

were provided in Section 17.2. These criteria should be reviewed to avoid 

excessively high and unreasonable blow count requirements. When piles 

must penetrate into very hard geomaterials, weathered bedrock, or hard 

bedrock to satisfy any pile penetration depth requirements, predrilled holes 

may be required as it may not be possible to achieve the required penetration 

depth through driving alone.

• Pile termination and acceptance criteria in soft and hard rock.
439

Oftentimes piles can be driven into soft rock at considerable distances with a 

gradual buildup of both the pile penetration resistance as well as the nominal 

geotechnical resistance. In general, this condition is not too problematic 

unless the driving occurs at high blow counts for an extended duration of 

driving resulting in a reduction in hammer performance. In the case of hard 

rock, driving should be terminated relatively quickly once hard rock is 

encountered to reduce the risk of pile toe or hammer damage as discussed in 

Section 17.2.

• Understanding the limitations of dynamic tests and wave equation analyses in 

easy driving (less than 24 blows per foot) and hard driving (more than 120 blows 

per foot) situations.

At pile penetration resistances less than 24 blows per foot and above 120 

blows per foot, dynamic test and analysis methods can overpredict and 

underpredict the nominal resistance, respectively. At low blow counts (high 

set per blow), it is difficult for dynamic methods to easily separate the static 

and dynamic soil resistance effects resulting in a tendency to overpredict the 

static resistance. Use of a reduced hammer stroke or lower fuel setting can 

help improve the accuracy of dynamic methods in low blow count situations. 

At very high blow counts (low set per blow), dynamic test methods tend to 

produce lower bound nominal resistance estimates as not all of the resistance 

(particularly at and near the toe) is fully activated. In these high blow count 

situations, use of a larger hammer stroke, higher fuel setting, pile hammer 

with a greater rated energy, or variable stroke drop hammer can help improve 

dynamic method accuracy.

• Understanding that static load testing, dynamic testing with signal matching, 

wave equation analysis results, and dynamic formula results will give different 

values of the nominal resistance on the same pile.

All of the above methods for nominal resistance verification have a different 

resistance factor, indicating they all differ in reliability. Therefore, it should be 

surprising if an identical nominal resistance were determined from each 

method. Driving criteria should be developed using the most reliable of the 

resistance verification methods available provided that the results from that 

resistance verification method are checked and appear reasonable.

In the case of a static load test, the correlating data is only the pile 

penetration depth, blow count at end of driving, and the associated hammer 
440

stroke. Other tools must therefore be used to determine what other stroke 

and blow count combinations (the load test stroke and blow count values) are 

suitable for the required nominal resistance at the same penetration depth 

(accomplished by refined wave equation analysis) or if the same blow count 

and stroke are suitable for the nominal resistance at other locations and pile 

penetration depths (accomplished by a review of subsurface conditions and 

static analysis). A procedure for calibrating wave equation generated driving 

criteria to the dynamic testing signal matching results was described in 

Section 12.6.9. A similar approach may be used to calibrate wave equation 

based driving criteria to the static load test as described in Section 17.4. 

• An understanding of how to perform and evaluate restrike dynamic test results.

Ideally, the hammer stroke or fuel setting is selected such that the penetration 

resistance at the beginning of restrike falls between 2 to 3 and 10 blows per 

inch. In this situation, the test record to select for signal matching analysis is 

readily apparent. An early, high energy blow, with good data quality should 

be selected and analyzed for the nominal resistance. The restrike blow count 

should be carefully recorded over the full restrike event as the rate by which 

the blow count decreases from inch to inch can be helpful. When the restrike 

blow count is near or less than 24 blows per foot, a lower energy restrike blow 

should be chosen for signal matching analysis to reduce the potential for 

overpredicting the nominal resistance.

In more difficult situations, limited pile movement may occur during restrike 

and several records may need to be analyzed with signal matching. 

Superposition of the activated shaft resistances under various restrike 

hammer blows may be used to assess the nominal geotechnical resistance. 

Initial restrike blows may mobilize the shaft resistance along the upper portion 

of the pile shaft. Later restrike blows may indicate more shaft resistance on 

the lower portion of the pile once the upper shaft resistance has started to 

breakdown. The toe resistance and shaft resistance on the lower portion of 

the pile from the end of drive analysis should also be reviewed and, if 

appropriate, used in a superposition case. When using the toe resistance 

from an end of drive situation, the analyst should be confident that relaxation 

in the bearing layer is not a consideration or overestimation of the nominal 

resistance could result by using superposition.

• Failure to make timely decision regarding the acceptance of production piles 

during installation. 
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The acceptability of production piles must be determined in a timely manner 

following completion of driving. Delay claims can result if decisions to splice 

and drive deeper, the need to add piles, or the need to drive replacement 

piles, cannot be made in a reasonable amount of time.

• Difficulty to accept that, in some cases, dynamic methods of estimating nominal 

resistance (dynamic formulas, wave equation and dynamic measurements) may 

yield conservative predictions of the true geotechnical resistance and thus, 

correlations and extrapolation between dynamic and static load test results are 

necessary.

Dynamic methods can yield conservative estimates of the true geotechnical 

resistance in some situations. For example, an undersized hammer will not 

be able to mobilize the full soil resistance. Occasionally also, open ended 

pipe piles or H-piles which do not bear on rock may behave differently under 

dynamic and static loading conditions. Under dynamic loading conditions, the 

soil inside a pipe pile or between H-pile flanges may slip and produce internal 

shaft resistances. Under static loading conditions, this soil may plug and 

move with the pile resulting in toe resistance over the full pile cross section. 

Hence both shaft and toe resistances may be different in open profile pile 

sections under static and dynamic loading conditions. Plugging behavior can 

also vary in different geomaterials. Careful interpretation and extrapolation of 

dynamic results is required in these situations. Additional commentary on pile 

driving of open pile sections is provided in Section 7.10.7. 

17.4 EXAMPLES FOR ESTABLISHING DRIVING CRITERIA

The following simplified examples illustrate the considerations necessary to establish 

both economical and safe installation criteria (a) for 2% dynamic testing or (b) for 

one static test plus 2% dynamic testing. Various other ways of determining a driving 

criterion could be envisioned. For example, based on initial testing results, a refined 

wave equation bearing graph could be established, providing required blow counts 

for different nominal resistance values. Another example would be performing 

restrike dynamic testing after all piles had been installed to a wave equation 

calculated required blow count. That approach may be satisfactory if the bearing 

layer would be well documented and competent. However, if the final restrike tests 

would indicate insufficient nominal resistance, the piles would have to be redriven to 

a greater depth, or additional piles would be necessary, or 100% testing may be 
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necessary. The latter remedy would then allow for an increased resistance factor 

and, hopefully make the reduced nominal resistance values acceptable.

17.4.1 Driving Criterion – Example 1

For a small bridge project with less than 100 piles, the foundation designer 

determined that dynamic testing with signal matching should be used for the nominal 

resistance verification method and driving criteria. Therefore dynamic testing with 

signal matching was specified on 2% of the production piles using the AASHTO 

resistance factor, φdyn, of 0.65. The bridge project involved pile driving at three pier 

and two abutment locations. Soil borings indicated a relatively uniform site 

consisting of sandy soils.

All dynamic tests were performed prior to production pile driving began at each 

substructure location. One pile at each abutment or pier was tested for a total of five 

test piles. This quantity exceeded the requirement of a minimum of two test piles for 

the uniform site condition. The factored loads were 117 kips at the abutments and 

150 kips at the piers. This meant the required nominal resistance per pile was 

117kips / 0.65 or 180 kips at the abutments and 150 kips / 0.65 or 230 kips at the 

piers. Piles were dynamically tested during initial driving and again during restrike 

the next day as applicable for the sandy soil conditions. Signal matching was

performed on the dynamic test data acquired at the end-of-driving and beginning of 

restrike for each of the five test piles. A summary of the test results is presented in 

Table 17-1. 

Table 17-1 Example 1 Summary of Requirements and Results

Test Pile 

Location

Factored 

Load

(kips) 

Required 

Nominal 

Resistance 

(kips) 

Test

Pile 

Blow 

Count 

(bl/ft)

Nominal 

Resistance 

from Signal 

Matching 

at EOD 

(kips) 

Nominal 

Resistance 

from Signal 

Matching 

at BOR 

(kips) 

Soil Setup 

Factor

N Abut. 117 180 40 185 215 1.16

Pier 1 150 230 47 235 275 1.17

Pier 2 150 230 49 240 290 1.21

Pier 3 150 230 50 245 290 1.18

S Abut. 117 180 42 188 215 1.14

At the abutments, the average test pile nominal resistances from signal matching at 

the end of driving and restrike were 187 and 215 kips respectively; the average 
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setup factor was, therefore, 1.15. The average driving resistance at the end of 

driving varied between 40 and 42 blows/ft.

Using the nominal resistances calculated by signal matching, it was conservatively 

decided that for the production piles, the EOD nominal resistances could be 10% 

less than the required nominal resistance of 180 kips. The associated required 

penetration resistance, i.e., the driving criterion, was determined through refined 

wave equation analysis to be 36 blows/ft. This yielded a 164 kip nominal resistance 

at end of driving which with setup provided the 180 kip nominal resistance.

At the piers, the average test pile capacity from signal matching at the end of driving 

and restrike were 240 and 285 kips respectively; the average setup factor was, 

therefore, 1.19. The average driving resistance at the end of driving varied between 

47 and 50 blows/ft.

Using the nominal resistances calculated by signal matching, it was once again 

conservatively decided that for the production piles, the EOD nominal resistances 

could be 10% less than the required nominal resistance of 230 kips. The associated 

required penetration resistance was determined through refined wave equation 

analysis to be 43 blows/ft. This yielded a 209 kip nominal resistance at end of driving 

which with setup provided the 230 kip nominal resistance.

Note: The test piles were driven to nominal resistance values in excess of those 

required, and were, therefore, acceptable as production piles.

17.4.2 Driving Criterion – Example 2

For a relatively small bridge project with two abutments, three piers and less than 

120 piles, the foundation designer determined that static load testing and dynamic 

testing of 2% of the production piles (AASHTO resistance factor φdyn, of 0.80) should 

be used for the nominal resistance verification method and driving criteria. 

Therefore a static load test was planned at one pier along with dynamic testing and 

signal matching on 2% of the production piles. The bridge project involved pile 

driving at three pier and 2 abutment locations. Subsurface conditions are relatively 

uniform across the bridge site and consist primarily of medium dense fine sands.

Based on factored loads of 144 kips at the abutments and 184 kips at the piers, the 

required nominal resistances per pile were 180 kips at the abutments and 230 kips 

at the piers. The static load test frame and reaction system was set up at a pier 
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location for a target loading capacity of 230 kips, but with a maximum test load 

capability of 350 kips.

The static test loading was preceded by dynamically monitoring the installation of 

first the reaction piles and then the static load test pile. The reaction piles were 

driven to a penetration depth determined by static analysis as required for their uplift 

load (350 kips/number of reaction piles) and to a corresponding driving resistance 

determined by wave equation analysis. The static load test pile was driven to a 

depth determined by static analysis for the required 230-kip nominal resistance and 

this was achieved at a final penetration resistance of 60 blows/ft. The load test 

frame was then constructed and one week after installation, the static load test was 

performed. This was followed by dynamic restrike tests of the static load test pile 

and all reaction piles.

The signal matching result at the end of driving on the load test pile indicated a 

nominal resistance of 220 kips. The static load test evaluated by the Davisson 

Criterion, failed at a nominal resistance of 260 kips. Signal matching on the restrike 

dynamic test data indicated a nominal resistance of 290 kips. While this 30 kip 

difference in nominal resistance between the dynamic test restrike and static load 

test was likely caused by the preloading of the granular soils by the static load test 

performed before the restrike, the geotechnical engineer attributed it to a systemic 

difference between the test methods. For conservatism, a correlation factor of 260 

kips / 290 kips or 0.9 was recommended to be applied to all dynamic test results. 

For the load test pile this meant that the agreed upon EOD nominal resistance was 

90% of 220 kips or 198 kips which yielded a soil setup factor of 260 kips / 198 kips 

or 1.31. The geotechnical engineer also recommended a conservative soil setup 

factor of 1.2 be used across the relatively uniform site which then required that the 

piles be driven to EOD nominal resistances confirmed by signal matching of 167 kips 

(180 kips / (1.2* 0.9) = 167 kips) at the abutments and 212 kips (230 kips / (1.2*0.9) 

= 212 kips) at the piers.

One production pile in each pier or abutment location was then dynamically tested 

and analyzed in this manner. When the test pile reached the required EOD nominal 

resistance, the associated penetration resistance or blow count was established as 

the driving criterion for the remainder of the production piles at that pier or abutment.
445

REFERENCES

American Association of State Highway and Transportation Officials (AASHTO). 

(2014). AASHTO LRFD Bridge Design Specifications, US Customary Units, 

Seventh Edition, with 2015 Interim Revisions. American Association of State 

Highway and Transportation Officials, Washington, D.C., 1960 p.

Brown, D.A., and Thompson, W.R. (2011). Developing Production Pile Criteria from 

Test Pile Data. National Cooperative Highway Research Program (NCHRP) 

Synthesis 418, Washington, D.C., 54 p.
446
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CHAPTER 18

CONSTRUCTION MONITORING OF PILE INSTALLATION

Knowledgeable construction monitoring and inspection play a critical role in the 

proper installation of pile foundations. The general trend in cost effective pile 

foundation design is to use fewer piles each with a higher nominal resistance. This 

often requires the use of larger pile sections and bigger installation equipment. The 

construction monitoring of these pile installations becomes critical because of less 

redundancy (fewer piles required), smaller tolerances, and higher resistance factors.

Construction monitoring is only as good as the knowledge, experience and 

qualifications of the “inspector”. The role and duties of inspection personnel will vary 

depending upon the contract delivery method. However, pile inspection is still 

required for foundation acceptance regardless of how the necessary tasks are 

contractually assigned. Inspection personnel must understand their role on the 

project as well as the operation of the hammer and its accessories, the pile behavior, 

the soil conditions, and how these components interact. Most pile installation 

problems are avoidable if the inspector uses systematic inspection procedures 

coupled with good communication and cooperation with the contractor.

The inspector must be more than just a "blow counter". The inspector is the "eyes 

and ears" for the owner and engineer. Timely observations, suggestions, reporting, 

and correction advice can ultimately assure the success of the project. The earlier a 

problem or unusual condition is detected and reported by the inspector, the earlier a 

solution or correction in procedures can be applied, and hence a potentially negative 

situation can be limited to a manageable size. If the same problem is left 

unattended, the number of piles affected increases, as do the cost of remediation 

and the potential for claims or project delays. Thus, early detection and reporting of 

any problem may be critical to keep the project on schedule and within budget.

An outline of construction monitoring procedures and maintenance of pile driving 

records is provided in this chapter. Further details on the inspection of piles and pile 

driving systems may be found in the Deep Foundations Institute (1995); (1997), 

provided at the end of the chapter. The FHWA document “Performance of Pile 

Driving Systems” by Rausche et al. (1986), as well as hammer and equipment 

manufacturer’s literature are also good sources for information and details on 
448

hammer operation. The Florida Department of Transportation also developed and 

continues to update an excellent inspection manual for driven pile foundations as 

part of their inspector training program, Passe (1994). The pile installation 

monitoring procedures and record keeping methods presented herein should be 

refined as needed based on the project delivery method and agency practice.

18.1 MONITORING NEEDS BASED ON THE PROJECT DELIVERY METHOD

Transportation projects have historically been designed and constructed using the 

design-bid-build (DBB) delivery method. Design-Build (DB) and Construction 

Manager / General Contractor (CMGC) project delivery methods are increasingly 

being used due to the reduced project development and delivery time associated 

with these methods. An overview of the key construction monitoring items on driven 

pile foundations projects and how the typical duties associated with that item vary 

depending on DBB or DB contract delivery is summarized in Table 18-1. 

Construction monitoring duties for the driven pile foundations items noted in Table 

18-1 for CMGC project delivery contracts may fall under either DBB or DB 

depending on the individual project.

18.2 ITEMS TO BE MONITORED

Regardless of the project delivery method there are several items to be monitored by 

the “inspector” on every pile foundation project. These include test piles driven to 

establish order lengths or for load testing, as well as for production piles. Items to 

be inspected can be grouped under one of the following areas:

1. Review of the foundation design report, project plans and specifications prior 

to the arrival at the project site.

2. Inspection of piles prior to installation.

3. Review contractor’s pile installation plan.

4. Inspection of pile driving equipment both before and during operation.

5. Inspection of test or indicator piles installation.

6. Inspection during production pile driving and maintenance of driving records.
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Table 18-1 Overview of Key Construction Monitoring and Inspection Items and 

Agency Inspection Duties on Design-Bid-Build and Design Build Contracts. 

Item Design-Bid-Build Design-Build 

Pile Installation 

Plan including 

Hammer 

Submittal 

Review pile installation plan. 

Perform or review required 

analyses for hammer approval if 

hammer approval performed.

Submit DB team pile installation 

plan to Engineer for conformance 

check with contract documents 

and requirements.

Test Pile 

Installation 

Inspect piles, hammer and 

appurtenances before driving. 

Inspect and record test pile 

driving, pile splices, pile 

alignment, location, and 

reference elevation.

Observe test pile installation and 

documentation by others. 

Communicate any concerns to 

Engineer.

Nominal 

Resistance 

Verification 

Perform resistance verification 

observations (blow count and 

stroke) or tests (static, dynamic). 

Review results of resistance 

verification method test.

Driving Criteria 

and Production 

Pile Order 

Lengths

Analyze test pile and resistance 

verification results. Establish 

driving criteria and determine 

order lengths.

Review driving criteria 

established by design-build team. 

Production Pile 

Installation 

Inspect piles, inspect pile splices, 

inspect and record production 

pile driving, document final 

alignment, location, and 

elevations.

Check all piles met the 

established driving criteria and 

associated plan requirements.

Foundation 

Certification 

Required documentation 

completed as part of overall 

inspection process.

Review foundation certification 

package from design-build team 

with Engineer.

Verification 

Testing Not applicable. Select piles for verification testing in coordination with Engineer.

Piling Problems 

and Resolution

Identify, document, and evaluate 

any piling problems. Coordinate 

problem resolution with Engineer, 

Designer, and Contractor.

Document resolution of any noted 

deficiencies. 

A flow chart identifying the key components of the pile inspection process is 

presented in Figure 18-1. On DBB projects the Contact Engineer corresponds to the 

agency engineer. On DB projects, the Contact Engineer is the design-build team 

foundation engineer of record.
450

Figure 18-1 Key components of the pile installation inspection process.
451

18.3 REVIEW OF PROJECT PLANS AND SPECIFICATIONS

The first task in construction monitoring or inspection is to thoroughly review the 

project plans and specifications as they pertain to pile foundations. All equipment 

and procedures specified, including any indicator or test program of static and/or 

dynamic testing, should be clearly understood. If questions arise, clarification should 

be obtained from the originator of the specifications. The preliminary driving criteria 

should be known, as well as methods for using the test program results to adjust the 

criteria to site specific hammer performance and soil conditions. The pile inspector 

should fully understand the responsibility of his/her organization in the DB or DBB 

project and should have answers to the following questions:

1. Is the inspector on the project in an observational capacity reporting to the 

foundation designer?, or

2. Does his/her organization have the direct responsibility to make decisions 

during driving of the test pile(s) and/or the production piles?

The inspector should also know:

1. Whom to contact if something goes wrong, and/or where to seek advice.

2. Whom to send copies of driving records and daily inspection reports.

3. What is required in the construction monitoring reports during pile driving 

activities and upon completion of the project.

4. How to inform contractor when work deviates from contract documents 

without directing contractor’s work.

18.4 INSPECTORS TOOLS

The checklist shown in Figure 18-2 is modified from Williams Earth Science (1995) 

and summarizes the tools a pile inspector should have readily available to perform 

their job.
452

Figure 18-2 Key components of the pile installation process 

(modified from Williams Earth Science 1995). 

18.5 INSPECTION OF PILES PRIOR TO AND DURING INSTALLATION

The inspection check list will be different for each type of pile, but some items will be 

the same. A certificate of compliance for the piles is generally required by the 

specifications. The inspector should obtain this certificate from the contractor and 

compare the specification requirements with the information provided on the 

certificate. The following sections contain specific guidance for each major pile type. 

Section 18.7 provides similar sections for each major hammer type. A detailed pile 

driving inspection list for a project can be obtained by combining the check list for 

that projects pile type in Section 18.5, and hammer type in Section 18.7.

18.5.1 Timber Piles

Physical details for round timber piles are sometimes referred to in the ASTM pile 

specification, ASTM D25. Regardless of the referenced specifications, the following 

items should be checked for compliance:
453

a. The timber should be of the specified species.

b. The piles should have the specified minimum length, and have the correct pile 

toe and butt sizes. The pile butt must be cut squarely with the pile axis.

c. The twist of spiral grain and the number and distribution of knots should be 

acceptable.

d. The piles should be acceptably straight.

e. The piles must be pressure treated as specified.

f. The pile butts and/or toe may require banding as detailed in Chapter 16.

g. Steel shoes which may be specified must be properly attached. Details are 

provided in Chapter 16.

h. Pile splices, if allowed by plans and specifications, must meet the project 

requirements.

18.5.2 Precast Concrete Piles

On many projects, inspection and supervision of casting operations for precast 

concrete piles is provided by the transportation agency. Frequently, in lieu of this 

inspection, a certificate of compliance is required from the contractor. The following 

checklist provides items to be inspected at the casting yard (when applicable):

a. Geometry and other characteristics of the forms.

b. Dimensions, quantity, and quality of spiral reinforcing and prestressing steel 

strands, including a certificate indicating that the prestressing steel meets 

specifications.

c. If the pile is to have mechanical or welded splices, or embedded toe 

protection, the splice or toe protection connection details including number, 

size and lengths of dowel bars should be checked for compliance with the 

approved details and for the required alignment tolerance. They should be 

cast within tolerance of the true axial alignment.

d. Quality of the concrete (mix, slump, strength, etc.) and curing conditions.
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e. Prestressing forces and procedures, including time of tension release, which 

is related to concrete strength at time of transfer.

f. Handling and storage procedures, including minimum curing time for concrete 

strength before removal of piles from forms.

The following is a list of items for prestressed concrete piles to be inspected at the 

construction site:

a. The piles should be of the specified length and section. Many specifications 

require a minimum waiting period after casting before driving is allowed. 

Alternatively, the inspector must be assured that a minimum concrete 

strength has been obtained. If the piles are to be spliced on the site, the 

splices should meet the specified requirements (type, alignment, etc.).

b. There should be no evidence that any pile has been damaged during shipping 

to the site, or during unloading of piles at the site. Lifting hooks are generally 

cast into the piling at pick up points. Piles should be unloaded by properly 

sized and tensioned slings attached to each lifting hook. Piles should be 

inspected for cracks or spalling.

c. The piles should be stored properly. When piles are being placed in storage, 

they should be stored above ground on adequate blocking in a manner which 

keeps them straight and prevents undue bending stresses.

d. The contractor should lift the piles into the leads properly and safely. Cables 

looped around the pile are satisfactory for lifting. Chain slings should never 

be permitted. Cables should be of sufficient strength and be in good 

condition. Frayed cables are unacceptable and should be replaced. For 

shorter piles, a single pick-up point may be acceptable. The pick-up point 

locations should be as specified by the casting yard. For longer piles, two or 

more pick up points at designated locations may be required.

e. The pile should be free to twist and move laterally in the helmet.

f. Piles should have no noticeable cracks when placed in leads or during 

installation. Spalling of the concrete at the top or near splices should not be 

evident.
455

18.5.3 Steel H-Piles

The following should be inspected at the construction site:

a. The piles should be of the specified steel grade, length, or section/weight.

b. Pile shoes, if required for pile toe protection, should be as specified. Pile 

shoe details are provided in Chapter 16.

c. Splices should be either proprietary splices or full penetration groove welds 

as specified. The top and bottom pile sections should be in good alignment 

before splicing. Pile splice details are discussed in Chapter 16.

d. Pile splices and pile toe attachments must be welded properly.

e. The piles being driven must be oriented with flanges in the correct direction 

as shown on the plans. Because the lateral resistance to bending of H-piles 

is considerably more in the direction perpendicular to flanges, the correct 

orientation of H-piles is very important.

f. There should be no observable pile damage, including deformations at the 

pile head.

18.5.4 Steel Pipe Piles

The following should be inspected at the construction site:

a. The piles should be of specified steel grade, length, and minimum 

section/weight (wall thickness) and either seamless or spiral welded as 

specified.

b. Piles should be driven either open ended or closed ended. Closed-ended 

pipe piles should have bottom closure plates or conical points of the correct 

size (diameter and thickness) and be welded on properly, as specified. Open 

end pipe piles should have cutting shoes that are welded on properly.

c. The top and bottom pile sections should be in good alignment before splicing. 

Splices or full penetration groove welds should be installed as specified. Pile 

splice details are discussed in Chapter 16.
456

d. There should be no observable pile damage, including deformations at the 

pile head. After installation, closed-end pipes should be visually inspected for 

damage or water prior to filling with concrete.

18.6 INSPECTION OF DRIVING EQUIPMENT

A typical driving system consists of crane, leads, hammer, hammer cushion, helmet, 

and in the case of concrete piles, a pile cushion. As discussed in Chapter 15, each 

component of the drive system has a specific function and plays an important role in 

the pile installation. The project plans and specifications may specify or restrict 

certain items of driving equipment. The inspector must check the contractor's driving 

equipment and obtain necessary information to determine conformity with the plans 

and specifications prior to the commencement of installation operations.

The following checklist will be useful in the inspection of pile driving equipment 

before driving:

1. The pile hammer should be the approved make and model as submitted or 

should meet specification requirements if no submittal is required. 

Usually the specifications require certain hammer types and/or specify minimum 

and/or maximum energy ratings. The inspector should make sure for single acting 

air/steam hammers that the contractor uses the proper size external power source 

and that, for adjustable stroke hammers, the stroke necessary for the required 

energy be obtained. For double acting or differential air/steam, the contractor must 

again obtain the proper size external power source and the operating pressure and 

volume must meet the hammer manufacturer's specification. For open end diesel 

hammers, the inspector should obtain a chart for determining stroke from visual 

observation, or alternatively have available a device for electronically estimating the 

stroke from the blow rate. For closed end diesel hammers, the contractor should 

supply the inspector with a calibration certificate for the bounce chamber pressure 

gauge and a chart which correlates the bounce chamber pressure with the energy 

developed by the hammer. The bounce chamber pressure gauge should be 

provided by the contractor. For single acting and double acting hydraulic hammers, 

the contractor should supply a system for measuring and displaying the hammer 

energy or impact velocity.
457

2. The hammer cushion being used should be checked to confirm it is of the 

approved material type, size and thickness.

The main function of the hammer cushion is to protect the hammer itself from fatigue 

and high frequency accelerations which would result from steel to steel impact with 

the helmet and/or pile. The hammer cushion should have the proper material and 

same shape/area to snugly fit inside the helmet (drive cap). If the cushion diameter 

is too small, the cushion will break or badly deform during hammer blows and 

become ineffective. The hammer cushion must not be excessively deformed or 

compressed. Some air/steam hammers rely upon a certain total thickness (of 

cushion plus striker plate) for proper valve timing. Hammers with incorrect hammer 

cushion thickness may not operate, or will have improper kinetic energy at impact. 

Since it is difficult to inspect this item once the driving operation begins, it should be 

checked before the contractor starts pile driving on a project as well as periodically 

during production driving on larger projects. A photograph of a hammer cushion 

ready for inspection prior to insertion into the helmet is presented in Figure 18-3. The 

Blue Nylon hammer cushion disks are shown in the lower right corner of the 

photograph. The hammer cushion thickness and diameter, the diameter of the 

helmet cushion pot, as well as the dimensions of the striker plate should all be 

measured by the inspector during a hammer cushion inspection. A damaged 

aluminum plate, found during a cushion check of an aluminum and micarta hammer 

cushion, is displayed on the left hand side of Figure 18-4. 

3. The helmet (drive cap) should properly fit the pile.

The purpose of the helmet is to hold the pile head in alignment and transfer the 

impact concentrically from the hammer to the pile. The helmet also houses the 

hammer cushion, and must accommodate the pile cushion thickness for concrete 

piles. The helmet should fit loosely to avoid transmission of torsion or bending 

forces, but not so loosely as to prevent the proper alignment of hammer and pile. 

Helmets should ideally be of roughly similar size to the pile diameter. Although 

generally discouraged, spacers may be used to adapt an oversize helmet, provided 

the pile will still be held concentrically with the hammer. A properly fitting helmet is 

important for all pile types, but is particularly critical for precast concrete piles. A 

poorly fitting helmet often results in pile head damage. Check and record the helmet 

weight for conformance to wave equation analysis or for future wave equation 

analysis. Larger weights will reduce the energy transfer to the pile.
458

Figure 18-3 Check of Blue Nylon hammer cushion material before use.

Figure 18-4 Damaged aluminum and micarta hammer cushion.
459

4. The pile cushion should be of correct type material and thickness for concrete 

piles.

The purpose of the pile cushion is to reduce high compression stresses, to evenly 

distribute the applied forces to protect the concrete pile head from damage, and to 

reduce the tension stresses in easy driving. Pile cushions for concrete piles should 

have the required thickness determined from a wave equation analysis but not less 

than 4 inches. A new plywood, hardwood, or composite wood pile cushion, which is 

not water soaked, should be used for every pile. In Figure 18-5, a new, 22 inch thick 

plywood pile cushion is being inserted into a helmet. The helmet with pile cushion 

inserted is shown in Figure 18-6. Note that after cushion insertion, minimal depth 

remains in the helmet to accommodate and adequately restrain the pile head.

Figure 18-5 New pile cushion being inserted into helmet. 
460

Figure 18-6 Helmet with new pile cushion installed.

The cushion material should be checked periodically for damage and replaced 

before excessive compression (more than half the original thickness), burning, or 

charring occurs. Wood cushions may take only about 1,000 to 2,000 blows before 

they deteriorate. During hard driving, more than one cushion may be necessary for a 

single pile. Longer piles or piles driven with larger hammers may require thicker pile 

cushions.

5. Predrilling, jetting or spudding equipment, if specified or permitted, should be 

available for use and meet the requirements.
461

The depth of predrilling, jetting or spudding should be very carefully controlled so 

that it does not exceed the allowable limits. Predrilling, jetting, or spudding below 

the allowed depths will generally result in a reduced nominal geotechnical 

resistance, and the pile acceptance may become questionable. Additional details on 

predrilling, jetting, and spudding equipment are presented in Chapter 15.

6. The lead system being used must conform to the requirements, if any, in the 

specifications. Lead system details are discussed in detail in Chapter 15.

The leads perform the very important function of holding the hammer and pile in 

good alignment with each other. Poor alignment reduces energy transfer as some 

energy is then imparted into horizontal motion. Poor alignment also generally results 

in higher bending stresses and higher local contact stresses which can cause pile 

damage. This is particularly important at end of driving when blow counts are 

highest and driving stresses are generally increased. Sometimes the specifications 

do not allow certain lead systems or may require a certain type system. A pile gate 

at the lead bottom which properly centers the pile should be required, as it helps 

maintain good alignment.

Note: On many projects, a wave equation analysis is used to determine preliminary 

driving criteria for design and/or construction control. The contractor is 

usually required to provide a pile and driving equipment data form similar to 

Figure 15-50 and obtain prior approval from the agency. Even if wave 

equation analysis is not required, this form should be included in the project 

files so a wave equation analysis could be performed in the future. This form 

can also function as a check list for the inspector to compare the proposed 

equipment with the actual equipment on-site.

18.7 INSPECTION OF DRIVING EQUIPMENT DURING INSTALLATION

The main purpose of construction monitoring and inspection is to assure that piles 

are installed so that they meet the driving criteria and are undamaged. Driving 

criteria are often defined, in part, by a minimum pile penetration resistance or blow 

count that is measured in blows per foot or blows per inch. Driving criteria assure 

that piles have the required nominal resistance. The blow count, however, is 

dependent upon the performance of the pile driving hammer. The blow count will 

generally be lower when the hammer imparts higher energy and force to the pile, 

while the blow count will be higher if the hammer imparts lower energy and force to 

the pile. High blow counts can be due either to soil resistance or to a poorly 
462

performing hammer. Thus, the inspector must evaluate if the hammer is performing 

properly to assure that the driving criteria has been met and therefore the nominal 

resistance is achieved.

Each hammer has its own operating characteristics; the inspector should not blindly 

assume that the hammer on the project is in good working condition. Two different 

types of hammers with identical manufacturer’s rated energy will not drive the same 

pile in the same soil with the same blow count. In fact, two supposedly identical 

hammers (same make and model) may not have similar driving capability due to 

several factors including differing friction losses, valve timing, air supply hose typelength-condition, fuel type and intake amount, ring condition, and other maintenance 

status items. The inspector should become familiar with the proper operation of the 

hammer(s) used on site. The inspector may wish to contact the hammer 

manufacturer or supplier who generally will welcome the opportunity to supply 

further information. The inspector should review the operating characteristics for the 

hammer which are included in Chapter 15. The following checklists briefly 

summarize key hammer inspection issues.

18.7.1 Drop Hammers

a. Determine/confirm the ram weight. Ram weight can be calculated from the 

ram volume and steel density of 492 lb/ft3 if necessary.

b. The leads should have sufficient tolerance and/or the guides greased to allow 

the ram to fall without obstruction or binding.

c. Make sure the desired stroke is maintained. Low strokes will reduce energy. 

Excessively high strokes increase pile stresses and could cause pile damage.

d. Make sure the helmet stays properly seated on the pile and that the hammer 

and pile maintain alignment during operation.

e. Make sure the hammer hoist line is spooling out freely during the drop and at 

impact. If the hoist line drags, less energy will be delivered. If the crane 

operator catches the ram too early, not only is less energy delivered, but 

energy is transmitted into the hoist line, crane boom, and hoist, which could 

cause maintenance and/or safety problems.
463

18.7.2 Single Acting Air/Steam Hammers

a. Determine/confirm the ram weight. Ram weight can be calculated from the 

ram volume and steel density of 492 lb/ft3 if necessary. Check for and record 

any identifying labels as to hammer make, model and serial number.

b. Check the air or steam supply and confirm it is of adequate capacity to 

provide the required pressure and flow volume. Also check the number, 

length, diameter, and condition of the air/steam hoses. Manufacturers 

provide guidelines for proper compressors and supply hoses. Air should be 

blown through the hose before attaching it to the hammer. The motive fluid 

lubricator should occasionally be filled with the appropriate lubricant as 

specified by the manufacturer. During operation, check that the pressure at 

the compressor or boiler is equal to the rated pressure plus hose losses. The 

pressure should not vary significantly during driving. The photograph of an air 

compressor display panel in Figure 18-7 illustrates the discharge pressure 

dial that should be checked.

c. Visually inspect the slide bar and its cams for excessive wear. Some 

hammers can be equipped with a slide bar with dual set of cams to offer two 

different strokes. The stroke can be changed with a valve, usually operated 

from the ground. Measure the stroke being attained and confirm it meets 

specification.

d. Check that the columns or ram guides, piston rod, and slide bar are well 

greased.

e. For most air/steam hammers, the total thickness of hammer cushion and 

striker plate must match the hammer manufacturer's recommendation and the 

hammer cushion cavity in the helmet for proper valve timing and hammer 

operation. This thickness must be maintained and should be checked before 

placing the helmet into the leads, and thereafter by comparison of cam to 

valve position and/or gap between ram and hammer base when the ram is at 

rest on the pile top.
464

Figure 18-7 Air compressor display panel.

f. Make sure the helmet stays properly seated on the pile and that the hammer 

and pile maintain alignment during operation.

g. The ram and column keys used to fasten together hammer components 

should all be tight.

h. The hammer hoist line should always be slack, with the hammer's weight fully 

carried by the pile. Excessive tension in the hammer hoist line is a safety 

hazard and will reduce energy to the pile. Leads should always be used.

i. Compare the observed hammer speed in blows per minute near end of 

driving with the manufacturer's specifications. Blows per minute can be timed 

with a stopwatch or a saximeter. Slower operating rates may imply a short 

stroke (from inadequate pressure or volume, restricted or undersized hose, or 

inadequate lubrication) or improper valve timing (possibly from incorrect 

cushion thickness or worn parts). Erratic hammer operation, such as skipping 
465

blows, can result from improper cushion thickness, poor lubrication, foreign 

material in a valve, faulty valve/cam system, or loose hammer fasteners or 

keys.

j. As the blow count increases, the ram stroke may also increase, causing it to 

strike the upper hammer assembly and lifting the hammer ("racking") from the 

pile. If this behavior is detected, the air pressure flow should be reduced 

gradually until racking stops. The flow should not be overly restricted so that 

the stroke is reduced.

k. Some manufacturers void their warranty if the hammer is consistently 

operated above 10 blows per inch of penetration beyond short periods such 

as required when toe bearing piles are driven to rock. Therefore, in 

prolonged hard driving situations, it may be more desirable to use a larger 

hammer or stiffer pile section.

l. Common problems and problem indicators for air/steam hammers are 

summarized in Table 18-2. 

An inspection form for single and differential acting air/steam hammers is provided in 

Figure 18-8. The primary feature of this form is the three column area in the middle 

of the form. The left column illustrates the key objects of the driving system. The 

middle column contains the manufacturer's requirements for key objects and the 

right column is used to record the observed condition of those objects. This format 

allows the inspector to quickly identify potential problems and an immediate 

correction may be possible. The hammer inspection form is intended to be used 

periodically during the course of the project as a complement to the pile driving log.

The bottom portion of the hammer inspection form contains an area where 

observations at final driving should be recorded. This information may be 

particularly interesting to an engineer who has performed a wave equation analysis 

as the actual situation can then be compared to the analyzed one. Therefore, it is 

recommended that a copy of the completed hammer inspection form be provided to 

appropriate design and construction personnel.
466

Table 18-2 Common Problems and Indicators for Air/Steam Hammers 

(after Williams Earth Sciences 1995)

Common Problems Indicators

Air trip mechanism on hammer 

malfunctioning.

Erratic operation rates or air valve 

sticking open or close.

Cushion stack height not correct (affects 

timing of trip mechanism air valve). Erratic operation rates.

Compressor not supplying correct 

pressure and volume of air to hammer.

Blows per minute rate is varying either 

faster or slower than the manufacturer 

specified.

Air supply line kinked or tangled in leads, 

boom or other. Visually evident.

Moisture in air ices up hammer. Ice crystals exiting exhaust ports of 

hammer.

Lack of lubricant in air supply lines. Erratic operation rates.

Packing around air chest worn, allowing 

air blow by.

Ram raises slowly - blows per minute 

rate slower than manufacturer 

specifications - air leaking around piston 

shaft and air chest.

Nylon slide bar worn. Visually evident.

Ram columns not sufficiently greased. Visually evident.
467

Figure 18-8 Inspector’s form for single and differential acting air/steam hammers.
468

18.7.3 Double Acting or Differential Air/ Steam Hammers

a. Determine/confirm the ram weight. Ram weight can be calculated from the 

ram volume and steel density of 492 lbs/ft3 if necessary. Check for and 

record any identifying labels as to hammer make, model and serial number.

b. Check the air or steam supply and confirm it is of adequate capacity to 

provide the required pressure and flow volume. This is extremely important 

since approximately half the rated energy comes from the pressure on the 

ram during the downstroke. Check also the number, length, diameter, and 

condition of the air/steam hoses. Manufacturers provide guidelines for proper 

compressors and supply hoses. Air should be blown through the hose before 

attaching it to the hammer. The motive fluid lubricator should occasionally be 

filled with the appropriate lubricant as specified by the manufacturer. During 

operation, check that the pressure at the compressor or boiler is equal to the 

rated pressure plus hose losses. The pressure should not vary significantly 

during driving. Record the pressure at the beginning of driving.

c. Visually inspect the slide bar and its cams for excessive wear. Measure the 

stroke being attained and confirm that it meets specification.

d. Check that the columns or ram guides, piston rod, and slide bar are well 

greased.

e. For most air/steam hammers, the total thickness of hammer cushion and 

striker plate must match the hammer manufacturer's recommendation and the

hammer cushion cavity in the helmet for proper valve timing and hammer 

operation. This thickness must be maintained, and can be checked before 

assembly of the helmet into the leads, and thereafter by comparison of cam to 

valve position and/or gap between ram and hammer base when the ram is at 

rest on the pile.

f. Make sure the helmet stays properly seated on the pile and that the hammer 

and pile maintain alignment during operation.

g. The ram and column keys used to fasten together hammer components 

should all be tight.

h. The hammer hoist line should always be slack with the hammer's weight and 

be fully carried by the pile. Excessive tension in the hammer hoist line is a 
469

safety hazard and will reduce energy to the pile. Leads should always be 

used.

i. Compare the observed hammer speed in blows per minute near end of 

driving with the manufacturer's specifications. Blows per minute can be timed 

with a stopwatch or a saximeter. Slower operating rates may imply a short 

stroke (from inadequate pressure or volume, restricted or undersized hose, or 

inadequate lubrication) or improper valve timing (possibly from incorrect 

cushion thickness or worn parts). Erratic hammer operation, such as skipping 

blows, can result from improper cushion thickness, poor lubrication, foreign 

material in a valve, faulty valve/cam system, or loose hammer fasteners or 

keys. 

j. As the penetration resistance increases, the ram stroke may also increase, 

causing it to strike the upper hammer assembly and lifting the hammer 

(racking) from the pile. If this behavior is detected, the pressure flow should 

be reduced gradually until racking stops. This will result in a reduction in 

energy since the pressure also acts during the downstroke, thereby 

contributing to the rated energy. Record the final pressure. The flow should 

not be overly restricted so that the stroke is also reduced, causing a further 

reduction in energy. For optimum performance, the pressure flow should be 

kept as full as possible so that the hammer lift-off is imminent.

k. Some manufacturers void their warranty if the hammer is consistently 

operated above 10 blows per inch of penetration beyond short periods such 

as required when toe bearing piles are driven to rock. Therefore, in 

prolonged hard driving situations, it may be more desirable to use a larger 

hammer or stiffer pile section.

l. Record the final pressure and compare with manufacturer's energy rating at 

this pressure.

m. Common problems and problem indicators for air/steam hammers are 

summarized in Table 18-2. 

An inspection form for enclosed double acting air/steam hammers is provided in 

Figure 18-9. The primary feature of this form is the three column area in the middle 

of the form. The left column identifies key objects of the driving system. The middle 

column contains the manufacturer's requirements for key objects and the right 

column is used to record the observed condition of those objects. This format allows 
470

Figure 18-9 Inspector’s form for enclosed double acting air/steam hammers.
471

the inspector to quickly identify potential problems and an immediate correction may 

be possible. The hammer inspection form is intended to be used periodically during 

the course of a project as a complement to the pile driving log.

The bottom portion of the hammer inspection form contains an area where 

observations at final driving should be recorded. This information may be 

particularly interesting to an engineer who has performed a wave equation analysis 

as the actual situation can then be compared to the analyzed one. Therefore, it is 

recommended that a copy of the completed hammer inspection form be provided to 

appropriate design and construction personnel.

18.7.4 Single Acting Diesel Hammers

a. Determine/confirm that the hammer is the correct make and model. Check 

for and record any identifying labels as to hammer make, model and serial 

number.

b. Make sure all exhaust ports are open with all plugs removed.

c. Inspect the recoil dampener for condition and thickness. If this is excessively 

worn or of an improper thickness (consult manufacturer) it should be 

replaced. If the recoil dampener is too thin, the stroke will be reduced. 

Conversely, if it is too thick, or if cylinder does not rest on the dampener 

between blows, the ram could blow out the hammer top and become a safety 

hazard.

d. Check that lubrication of all grease nipples is regularly made. Most 

manufacturers recommend the impact block be greased every half hour of 

operation.

e. As the ram is visible between blows, check the ram for signs of uniform 

lubrication and ram rotation. Poor lubrication will increase friction and reduce 

energy to the pile.

f. Determine the hammer stroke, especially at end of driving or beginning of 

restrike. A "jump stick" attached to the cylinder is a safety hazard and should 

not be used. The stroke can be determined by a saximeter which measures 

the time between blows and then calculates the stroke. The ram stroke 
472

height, h, can also be calculated from this formula using the number of blows 

per minute (bpm) recorded:

 Eq. 18-1 

Where:

h = ram stroke (feet).

bpm = blow per minute (dimensionless).

The calculated stroke may require correction for batter or inclined piles. The 

inspector should always observe the ram rings and visually estimate the stroke using 

the manufacturer's chart.

g. As the blow count increases, the stroke should also increase. At the end of 

driving, if the ram fails to achieve the correct stroke (part of the driving criteria 

from a wave equation analysis), the cause could be lack of fuel. Most 

hammers have adjustable fuel pumps. Some have distinct fuel settings as 

shown in Figure 18-10(a), others are continuously variable as shown in Figure 

18-10(b), and some use a pressure pump as shown in Figure 18-11. Make 

sure the pump is on the correct fuel setting or pressure necessary to develop 

the required stroke. The fuel and fuel line should be free of dirt or other 

contaminants. A clogged or defective fuel injector will also reduce the stroke 

and should be replaced if needed.

h. Low strokes could be due to poor compression caused by worn or defective 

piston or anvil rings. Check compression by raising the ram, and with the fuel 

turned off, allowing the ram to fall. The ram should bounce several times if 

the piston and anvil rings are satisfactory.

i. Watch for signs of pre-ignition. When a hammer preignites, the fuel burns 

before impact, requiring extra energy to compress gas and leaving less 

energy to transfer to the pile. In long sustained periods of driving, or if the 

wrong fuel with a low flash point is used, the hammer could overheat and 

preignite. When pre-ignition occurs, less energy is transferred and the blow 

count rises, giving a false indication of high nominal resistance. If piles driven 

with a cold hammer drive deeper or with less hammer blows, or if the blow 

count decreases after short breaks, pre-ignition could be the cause and 

should be investigated. Dynamic testing is the preferable method to check for 

pre-ignition. 
473

Figure 18-10 Fuel Pumps: (a) fixed four step pump and (b) variable fuel pump.

Figure 18-11 Hydraulic pump for fuel pump adjustments.
474

j. For some diesel hammers, the total thickness of hammer cushion and striker 

plate must match the hammer manufacturer's recommendation and the 

hammer cushion cavity in the helmet for proper fuel injection and hammer 

operation. This total thickness must be maintained.

k. Make sure the helmet stays properly seated on the pile and that the hammer 

and pile maintain alignment during operation.

l. The hammer hoist line should always be slack, with the hammer's weight fully 

carried by the pile. Excessive tension in the hammer hoist line is a safety 

hazard and will reduce energy to the pile. Leads should always be used.

m. ....Some manufacturers void their warranty if the hammer is consistently 

operated above 10 blows per inch of penetration beyond short periods, such 

as those required when toe bearing piles are driven to rock. Therefore, in 

prolonged hard driving situations, it may be more desirable to use a larger 

hammer or stiffer pile section.

n. Common problems and problem indicators for single acting diesel hammers 

are presented in Table 18-3. 

An inspection form for single acting diesel hammers is provided in Figure 18-12. 

The primary feature of this form is the three column area in the middle of the form. 

The left column identifies key objects of the driving system, the middle column 

contains the manufacturer's requirements for that object and the right column is used 

to record the observed condition of that object. This format allows the inspector to 

quickly identify potential problems and an immediate correction may be possible. 

The hammer inspection form is intended to be used periodically during the course of 

a project as a complement to the pile driving log.

The bottom portion of the hammer inspection form contains an area where 

observations at final driving should be recorded. This information may be 

particularly interesting to an engineer who has performed a wave equation analysis 

as the actual situation can then be compared to the analyzed one. Therefore, it is 

recommended that a copy of the completed hammer inspection form be provided to 

appropriate design and construction personnel.
475

Figure 18-12 Inspector’s form for single acting diesel hammers.
476

Table 18-3 Common Problems and Indicators for Single Acting Diesel Hammers 

(after Williams Earth Sciences 1995)

Common Problems Indicators

Water in fuel. Hollow sound, white smoke.

Fuel lines clogged. No smoke or little gray smoke.

Fuel pump malfunctioning. Inconsistent ram strokes, little gray 

smoke or black smoke.

Fuel injectors malfunctioning. Inconsistent ram strokes, little gray 

smoke or black smoke.

Oil low. Blows per minute rate is lower than 

specified.

Oil pump malfunctioning. Blows per minute rate is lower than 

specified.

Water in combustion chamber. Hollow sound, white smoke.

Piston rings worn. Low strokes.

Tripping device broken.

Pawl or pin used to lift piston does not 

engage piston.

Pawl engages but does not lift piston.

Over heating. Paint and oil on cooling fins start to 

burn/sound changes.

18.7.5 Double Acting Diesel Hammers

a. Determine/confirm that the hammer is the correct make and model. Check 

for and record any identifying labels as to hammer make, model and serial 

number.

b. Make sure all exhaust ports are open with all plugs removed.

c. Inspect the recoil dampener for condition and thickness. If excessively worn 

or of improper thickness (consult manufacturer), it should be replaced. If it is 

too thin, the stroke will be reduced. If it is too thick or if cylinder does not rest 

on dampener between blows, the ram will cause hammer lift-off.

d. Check that lubrication of all grease nipples is regularly made. Most 

manufacturers recommend the impact block be greased every half hour of 

operation.
477

e. After the hammer is stopped, check the ram for signs of lubrication by looking 

into the exhaust port or trip slot. Poor lubrication increases friction, thus 

reducing energy to the pile.

f. Always measure the bounce chamber pressure, especially at end of driving or 

restrike. This indirectly measures the equivalent stroke or energy. All double 

acting diesels have a gauge. On most hammers an external gauge is 

connected by a hose to the bounce chamber. A photograph of a typical 

external bounce chamber pressure gauge is presented in Figure 18-13. The 

manufacturer should supply a chart relating the bounce chamber pressure for 

a specific hose size/length to the rated energy. The inspector should 

compare measured bounce chamber pressure with the manufacturer's chart 

to estimate the energy. The bounce chamber pressure measured may 

require correction for batter or inclined piles.

g. As the penetration resistance increases, the stroke and bounce chamber 

pressure should also increase. If the ram fails to achieve the correct stroke or 

bounce chamber pressure (part of the driving criteria from a wave equation 

analysis) at final driving, the cause could be lack of fuel. All these hammers 

have continuously variable fuel pumps. Check that the fuel pump is on the 

correct fuel setting. The fuel should be free of dirt or other contaminants. A 

clogged or defective fuel injector reduces the stroke.

h. In hard driving, high strokes cause high bounce chamber pressures. If the 

cylinder weight cannot balance the bounce chamber pressure, the hammer 

will lift-off of the pile and the operator must reduce the fuel to prevent this 

unstable racking behavior. Ideally it is set and maintained so that lift-off is 

imminent. The bounce chamber pressure gauge reading should correspond 

to the hammer's maximum bounce chamber pressure for the hose length 

used when lift-off is imminent. If not, then the bounce chamber pressure 

gauge is out of calibration and should be replaced, or the bounce chamber 

pressure tank needs to be drained.

i. Low strokes indicated by a low bounce chamber pressure could be due to 

poor compression caused by worn or defective piston or anvil rings. Check

compression with the fuel turned off by allowing the ram to fall. The ram 

should bounce several times if the piston and anvil rings are satisfactory.
478

Figure 18-13 Typical external bounce chamber pressure gauge.

j. Watch for pre-ignition. When a hammer preignites, the fuel burns before 

impact requiring extra energy to compress the gas and reducing energy 

transferred to the pile. When pre-ignition occurs, the blow count increases 

giving a false indication of high nominal resistance. In long sustained periods 

of driving or if low flash point fuel is used, the hammer could overheat and 

preignite. If piles driven with a cold hammer drive deeper or with fewer 

hammer blows, or if the blow count decreases after short breaks, investigate 

for pre-ignition, preferably with dynamic testing.

k. For some diesel hammers, the total thickness of the hammer cushion and 

striker plate must match the manufacturer's recommendation for proper fuel 

injection timing and hammer operation. This total thickness must be 

maintained.

l. Make sure the helmet stays properly seated on the pile and that the hammer 

and pile maintain alignment during operation.
479

m. The hammer hoist line should always be slack, with the hammer's weight fully 

carried by the pile. Excessive tension in the hammer hoist line is a safety 

hazard and will reduce energy to the pile. Leads should always be used.

n. Some manufacturers void their warranty if the hammer is consistently 

operated above 10 blows per inch of penetration beyond short periods such 

as those required when toe bearing piles are driven to rock. Therefore, in 

prolonged hard driving situations, it may be more desirable to use a larger 

hammer or stiffer pile section.

o. Common problems and problem indicators for double acting diesel hammers 

are presented in Table 18-4. 

An inspection form for double acting diesel hammers is provided in Figure 18-14. 

The primary feature of this form is the three column area in the middle of the form. 

The left column identifies key objects of the driving system, the middle column 

contains the manufacturer's requirements for that object and the right column is used 

to record the observed condition of that object. This format allows the inspector to 

quickly identify potential problems and an immediate correction may be possible. 

The hammer inspection form is intended to be used periodically during the course of 

a project as a complement to the pile driving log.

The bottom portion of the hammer inspection form contains an area where 

observations at final driving should be recorded. This information may be 

particularly interesting to an engineer who has performed a wave equation analysis 

as the actual situation can then be compared to the analyzed one. Therefore, it is 

recommended that a copy of the completed hammer inspection form be provided to 

appropriate design and construction personnel.
480

Table 18-4 Common Problems and Indicators for Double Acting Diesel Hammers 

(after Williams Earth Sciences 1995)

Common Problems Indicators

Water in fuel. Hollow sound, white smoke.

Fuel lines clogged. No smoke or little gray smoke.

Fuel pump malfunctioning. Inconsistent ram strokes, little gray 

smoke or black smoke.

Fuel injectors malfunctioning. Inconsistent ram strokes, little gray 

smoke or black smoke.

Oil low. Blows per minute rate is lower than 

specified.

Oil pump malfunctioning. Blows per minute rate is lower than 

specified.

Build-up of oil in bounce chamber. Not visible from exterior.

Water in combustion chamber. Hollow sound, white smoke.

Piston rings worn. Low strokes.

Tripping device broken.

Pawl or pin used to lift piston does not 

engage piston.

Pawl engages but does not lift piston.

Over heating. Paint and oil on cooling fins start to burn/ 

sound changes.
481

Figure 18-14 Inspector’s form for double acting diesel hammers.
482

18.7.6 Single Acting Hydraulic Hammers

a. Determine/confirm the ram weight. If necessary, the ram weight can be 

calculated from the ram volume and steel density of 492 lbs/ft3 although some 

rams may be hollow or filled with lead. There may also be identifying labels 

as to hammer make, model, and serial number which should be recorded.

b. Check the power supply and confirm it has adequate capacity to provide the 

required pressure and flow volume. Also, check the number, length, 

diameter, and condition of the hoses (no leaks in hoses or connections). 

Manufacturers provide guidelines for power supplies and supply hoses. 

Hoses bent to a radius less than recommended could adversely affect 

hammer operation or cause hose failure.

c. Hydraulic hammers must be kept clean and free from dirt and water. Check 

the hydraulic filter for blocked elements. Most units have a built in warning or 

diagnostic system.

d. Check that the hydraulic power supply is operating at the correct speed and 

pressure. Allow the hammer to warm up before operation, and do not turn off 

power pack immediately after driving.

e. For single acting hydraulic hammers with observable rams, measure the 

stroke being attained and confirm that it meets specification. For hammers 

with enclosed rams, it is impossible to observe the ram and estimate the 

stroke.

f. Check that the ram guides and piston rod are well greased.

g. Where applicable, the total thickness of hammer cushion and striker plate 

must be maintained to match the manufacturer's recommendation for proper 

valve timing and hammer operation.

h. Make sure the helmet stays properly seated on the pile and that the hammer 

and pile maintain alignment during operation.

i. The hammer hoist line should always be slack, with the hammer's weight fully 

carried by the pile. Excessive tension in the hammer hoist line is a safety 

hazard and will reduce energy to the pile. Leads should always be used.
483

j. Compare the observed hammer speed in blows per minute from near end of 

driving with the manufacturer's specifications. Blows per minute can be timed 

with a stopwatch or a saximeter. Slower operating rates at full stroke may 

imply excessive friction, or incorrect hydraulic power supply.

k. As the penetration resistance increases, the ram stroke may also increase, 

causing the ram to strike the upper hammer assembly and lifting the hammer 

from the pile (racking). If this behavior is detected, the pressure flow should 

be reduced gradually until racking stops. Many of these hammers have 

sensors, and if they detect this condition, the hammer will automatically shut 

down. The flow should not be overly restricted so that the correct stroke is 

maintained.

l. Some manufacturers void their warranty if the hammer is consistently 

operated above 10 blows per inch of penetration beyond short periods such 

as those required when toe bearing piles are driven to rock. Therefore, in 

prolonged hard driving situations, it may be more desirable to use a larger 

hammer or stiffer pile section.

m. Common problems and problem indicators for hydraulic hammers are 

summarized in Table 18-5. 

Table 18-5 Common Problems and Indicators for Single Acting Hydraulic 

Hammers (after Williams Earth Sciences 1995)

Common Problems Indicators

Hoses getting caught in leads. Visually evident.

Fittings leaking. Hydraulic fluid dripping.

Electrical connections. Erratic performance.

Sensors. Erratic performance.
484

Figure 18-15 Inspector’s form for single acting hydraulic hammers.
485

18.7.7 Double Acting Hydraulic Hammers

a. Determine/confirm the ram weight. There may also be identifying labels as to 

hammer make, model, and serial number which should be recorded.

b. Check the power supply and confirm it has adequate capacity to provide the 

required pressure and flow volume. Also, check the number, length, 

diameter, and condition of the hoses (no leaks in hoses or connections). 

Manufacturers provide guidelines for power supplies and supply hoses. 

Hoses bent to a radius less than recommended could adversely affect 

hammer operation or cause hose failure.

c. Hydraulic hammers must be kept clean and free from dirt and water. Check 

the hydraulic filter for blocked elements. Most units have a built in warning or 

diagnostic system.

d. Check that the hydraulic power supply is operating at the correct speed and 

pressure. Check and record the pre-charge pressures or accumulators. 

Allow the hammer to warm up before operation, and do not turn off power 

pack immediately after driving.

e. Most double acting hydraulic hammers have built in sensors to determine the 

ram velocity just prior to impact. This result may be converted to kinetic 

energy or equivalent stroke. The inspector should verify that the correct ram 

weight is entered in the hammer's "computer". This monitored velocity, 

stroke, or energy result should be constantly monitored and recorded. Some 

hammers have, or can be equipped with, a printout device to record that 

particular hammer's performance information with pile penetration depth 

and/or blow count. This is the most important hammer check that the 

inspector can and should make for these hammers. A photograph of a 

hydraulic hammer readout panel mounted on the power pack is presented in 

Figure 18-16. Hand held displays are also available on some hammers.

f. Most double acting hydraulic hammers are fully enclosed and therefore do not 

have observable rams. On these hammers it is impossible to measure the 

stroke being attained and confirm that it meets specification. Properly 

working energy readout devices are therefore mandatory for inspection.
486

Figure 18-16 IHC hydraulic hammer read-out panel. 

g. Most double acting hydraulic hammers are steel ram on steel anvil impact 

and do not use a striker plate or hammer cushion. One manufacturer uses an 

aluminum stack for a hammer cushion. Double acting hydraulic hammer 

models designed for use on concrete piles have a synthetic anvil block 

instead of a steel one. Document the anvil or cushioning mechanism, 

dimensions used in the double acting hammer model.

h. Make sure the helmet stays properly seated on the pile and that the hammer 

and pile maintain alignment during operation.

i. The hammer hoist line should always be slack, with the hammer's weight fully 

carried by the pile. Excessive tension in the hammer hoist line is a safety 

hazard and will reduce energy to the pile. Leads should always be used.

j. Compare the observed hammer speed in blows per minute from near end of 

driving with the manufacturer's specifications. Blows per minute can be timed 

with a stopwatch or a saximeter. Slower operating rates at full stroke may 

imply excessive friction, or incorrect hydraulic power supply.
487

k. As the penetration resistance increases, the ram stroke may also increase, 

causing the ram to strike the upper hammer assembly and lifting the hammer 

from the pile (racking). If this behavior is detected, the pressure flow should 

be reduced gradually until racking stops. Many of these hammers have 

sensors, and if they detect this condition, the hammer will automatically shut 

down. The flow should not be overly restricted so that the correct stroke is 

maintained.

l. Some manufacturers void their warranty if the hammer is consistently 

operated above 10 blows per inch of penetration beyond short periods such 

as those required when toe bearing piles are driven to rock. Therefore, in 

prolonged hard driving situations, it may be more desirable to use a larger 

hammer or stiffer pile section.

m. Common problems and problem indicators for double acting hydraulic 

hammers with nitrogen caps are summarized in Table 18-6. 

Table 18-6 Common Problems and Indicators for Double Acting Hydraulic 

Hammers 

Common Problems Indicators

Hydraulic hoses crossed. Jumping hydraulic hoses.

Incorrect pressure in accumulators. Jumping hydraulic hoses.

Cap pressure too high. Hammer jumping on pile.

Low oil supply. Erratic hammer operation.

Pressure or return valve malfunctioning. Erratic or no hammer operation.

Hammer inoperable. Control cable not connected.

Hammer inoperable. Bulb A or B not lit on control panel.

Hoses getting caught in leads. Visually evident.

Fittings leaking. Hydraulic fluid dripping.

Electrical connections. Erratic performance.

Sensors. Erratic performance.
488

Figure 18-17 Inspector’s form for hydraulic hammers. 
489

18.7.8 Resonant Hammers

a. Confirm that the hammer make and model meets specifications. There may 

also be identifying labels as to hammer make, model and serial number which 

should be recorded.

b. Check the power supply to confirm adequate capacity to provide the required 

pressure and flow volume. Check also the number, length, diameter, and 

condition of the hoses (no leaks in hoses or connections). Manufacturers 

provide guidelines for proper power supplies and supply hoses. Hoses bent 

to a smaller radius than recommended could affect hammer operation or 

cause hose failure.

c. Resonant hammers must be kept clean and free from dirt and water. Check 

the hydraulic filter for blocked elements. Most units have a built in warning or 

diagnostic system.

d. Check and record that the hydraulic power supply is operating at the correct 

speed and pressure. Allow the hammer to warm up before operation, and do 

not turn off the power pack immediately after driving. 

e. Record the resonant vibrating frequency.

f. Make sure the hydraulic clamps for attachment to the pile are in good working 

order and effective.

g. The hammer hoist line should always be slack enough to allow penetration 

with the hammer's weight primarily carried by the pile. Excessive tension in 

the hammer hoist line will retard penetration. If used for extraction, the hoist 

line should be tight at all times. Leads may or may not be used.

18.7.9 Vibratory Hammers

a. Confirm that the hammer make and model meets specifications. There may 

also be identifying labels as to hammer make, model and serial number which 

should be recorded.

b. Check the power supply to confirm adequate capacity to provide the required 

pressure and flow volume. Check also the number, length, diameter, and 

condition of the hoses (no leaks in hoses or connections). Manufacturers 
490

provide guidelines for proper power supplies and supply hoses. Hoses bent 

to a smaller radius than recommended could affect hammer operation or 

cause hose failure.

c. Vibratory hammers must be kept clean and free from dirt and water. Check 

the hydraulic filter for blocked elements. Most units have a built in warning or 

diagnostic system. 

d. Check and record that the hydraulic power supply is operating at the correct 

speed and pressure. Allow the hammer to warm up before operation, and do 

not turn off the power pack immediately after driving.

e. Record, if available, the vibrating frequency.

f. Make sure the hydraulic clamps for attachment to the pile are in good working 

order and effective.

g. The hammer hoist line should always be slack enough to allow penetration 

with the hammer's weight primarily carried by the pile. Excessive tension in 

the hammer hoist line will retard penetration. If used for extraction, the hoist 

line should be tight at all times. Leads are rarely used.

18.8 INSPECTION OF TEST OR INDICATOR PILES

Most pile foundation projects required verification of the foundation design and 

nominal resistance through the testing of some selected piles. The size of the 

foundation and relative costs of testing often dictate the type and number, if any, of 

verification tests performed. The inspector may be responsible for coordinating the 

test pile program with the contractor, other state personnel, and/or outside testing 

agencies.

Small foundations with few piles may be designed conservatively with low resistance 

factors and greater pile lengths. On these projects, test piles for verification testing 

are frequently not required. All piles are then production piles, and the entire pile

foundation is usually installed in one or two days. Information on the piles, 

hammers, and other observations are recorded by the inspector and appropriately 

passed on or filed. Inspection should be thorough as it is the only assurance of a 

good foundation. If any problems are observed, such as very low blow counts, 

refusal driving above scour depths, or excessive pile lengths, the problems and all 
491

pertinent observations must be reported quickly so that immediate corrective action 

can be taken. 

On most projects, some level of verification testing is specified. Small to mid-size 

projects may have only a single static test (Chapter 9) on one pile at a specific 

depth, or there may be a few dynamic test piles (Chapter 10). The dynamic tests 

may include either testing during driving to assess hammer performance and driving 

stresses, or testing during restrike to assess nominal resistance, or both. The static 

or dynamic tests should be performed and reviewed by personnel having 

appropriate knowledge of the test method and proper procedures. Generally, tests 

are done on some of the first piles driven to verify or adjust the driving criteria which 

will then be used for subsequent production piles. This further verification provides 

rational basis for changes to the driving criteria, if necessary, which should be 

applied to subsequent production pile driving.

On larger projects, multiple test piles distributed across the site are often required to 

verify or adjust the driving criteria as conditions warrant. The goal is to determine 

driving criteria which will lead to a safe and economical foundation. Such tests could 

be primarily done at one time at the beginning of the construction. For example, socalled indicator piles are driven in selected locations across the project site to 

establish order lengths for concrete piles. Selected piles are generally statically 

and/or dynamically tested. Alternatively, testing could be performed as the 

construction progresses with some test(s) establishing the driving criteria for piles in 

close proximity to the test pile(s), followed by production pile driving, and then 

repeating the process in stages across the site.

The test piles are often the most critical part of the foundation installation. The 

procedures and driving criteria established during this phase will be applied to all 

subsequent production piles. The largest savings are often found at this time. For 

example, test results may determine that the design pile length results in a greater 

nominal resistance than required and that the piles could be made substantially 

shorter. Alternatively, problems with the test piles are usually followed by the same 

problems with production piles. Since problems are in themselves costly, and if left 

unresolved may eventually escalate, determination of the best solution as quickly as 

possible should be accomplished. It is the inspector's responsibility to be observant 

and communicate significant observations precisely and in a timely manner to the 

appropriate agency, design, or construction personnel.
492

The answers to the following questions should be known before driving test piles. It 

is often beneficial to have a preconstruction pile driving meeting between the 

contractor, project administrator, and the inspector to clarify these items.

1. Who determines test pile locations?

2. Who determines the test pile driving criteria?

3. Who stops the driving when the driving criterion is met?

4. Who decides at what depth to stop the indicator/test piles?

5. Who checks cutoff elevations? 

6. Who checks for heave?

7. Who determines if static test and/or dynamic test results indicate an 

acceptable test pile?

8. Who determines if additional tests are required?

9. Who determines if modifications to procedures or equipment are required?

10. What documentation is required from test pile installations (test pile driving 

record, dynamic test results, static test report) and who produces what 

documentation?

11. Who produces the production pile driving criteria and how quickly will it be 

available?

12. Who has authority to allow production pile installation to begin?

13. When is the authorization to proceed to production pile driving given relative 

to test pile driving?

14. Can production piles (initial pile section or entire piles) be driven in advance 

of the production pile driving criteria and if so at whose risk?
493

18.9 INSPECTION OF PRODUCTION PILES

During the production pile driving operations, the inspector's function is to apply the 

knowledge gained from the test program to each and every production pile. Quality 

assurance measures for the pile quality and pile splices; hammer operation and 

cushion replacement; overall evaluation of pile integrity; procedures for completing 

the piles (e.g. filling pipe piles with concrete); and unusual or unexpected 

occurrences need to be addressed. Complete documentation for each and every 

pile must be obtained, and then passed on to the appropriate authorities in a timely 

manner.

The following items should be checked frequently (e.g. for each production pile):

1. Is the pile the specified type, size, length, and strength?

2. Is the pile installed in the correct location, within acceptable tolerances, and 

with the correct orientation?

3. Are splices, if applicable, made to specification?

4. Is pile toe protection required and properly attached?

5. Is the pile acceptably plumb?

6. Is the hammer working correctly?

7. Is the hammer cushion the correct type and thickness?

8. Is the pile cushion the correct type and thickness? Is it being replaced 

regularly?

9. Did the pile meet the driving criteria as expected? 

10. Did the pile have unusual driving conditions and therefore potential problems?

11. Is there any indication of pile heave?

12. Is the pile cutoff at the correct elevation?

13. Is there any visual damage?
494

14. Have all pipe piles been visually inspected prior to concrete filling? Has it 

been filled with the specified strength concrete? Were concrete samples 

taken?

15. Are piles which are to be filled with concrete, such as open ended pipes and 

prestressed concrete piles with center voids, being cleaned properly after 

driving is completed?

16. If there is any question about pile integrity, has the issue been resolved? Is 

the pile acceptable, or does it need remediation or replacement?

17. Is the documentation for this pile complete, including driving log? Has it been 

submitted on a timely basis to the appropriate authority.

Previous sections of this chapter provide material which relate to inspection of 

production piles and offer detailed answers to the questions raised above. Although 

the inspector has now had the experience of test pile installation, a few additional 

details and concerns are perhaps appropriate.

Counting the number of hammer blows per minute and comparing it to the 

manufacturer's specification will provide a good indication of whether or not the 

hammer is working properly. The stroke of the hammer for most single and double 

acting air/steam hammers can be observed. Check the stroke of a single acting 

diesel hammer with a saximeter or by computation from the blows per minute using 

Equation 18-1. Check and record the bounce chamber pressure for double acting 

diesel hammers. The stroke of most single acting hydraulic hammers can be 

observed. Record the energy from the built-in energy monitor in addition to hammer 

stroke for each pile. Double hydraulic hammers must have a built-in energy monitor, 

and this information should be recorded for each pile. The hammer inspection form 

presented earlier in this chapter should be completed for the hammer type used.

A hammer cushion of manufactured material usually lasts for many hours of pile 

driving, (as much as 200 hours for some manufactured materials) so it is usually 

sufficient to check before the pile driving begins and periodically thereafter. Pile 

cushions (usually made of plywood) need frequent changing because of excessive 

compression or charring and have a typical life of about 1000 to 2000 hammer 

blows. Pile cushions should preferably be replaced as soon as they compress to 

one half of the original thickness, or if they begin to burn. No changes to the pile 

cushion thickness should be permitted near final driving. The required pile 
495

penetration resistance or blow count for nominal resistance verification should only 

be determined following the first 100 blows after cushion replacement. A new pile 

cushion reduces energy transfer and therefore produces an inflated blow count 

compared to a used cushion.

Inspection of pile splices is important to assure pile integrity. Poorly made splices 

are a potential source of problems and possible pile damage during driving. In some 

cases damage may be detected from the blow count records. Dynamic pile testing 

can be useful in questionable cases.

Pile driving stresses should be kept within specified limits. If dynamic monitoring 

equipment was used during test pile driving, the developed driving criteria should 

keep driving stresses within specified limits. If periodic dynamic tests are made, a 

check that the driving stresses remain within the specified limits can be provided. 

Adjustments of the ram stroke for all hammer types may be necessary to avoid pile 

damage. For concrete piles, cushion thicknesses or driving procedures may need 

adjustment to control tension and compression stresses. If dynamic testing is not 

used, a wave equation analysis is essential to evaluate the anticipated driving 

stresses. 

Driving of piles at high blow counts, above 10 blows per inch, should be avoided by 

matching the driving system with the pile type, length and subsurface conditions. 

This should have been accomplished in the design phase by performing wave 

equation analysis. However, conditions can change across the project due to site 

variability.

All piles should be checked for damage after driving is completed. The driving 

records for all pile types can be compared with adjacent piles for unusual records or 

vastly different penetrations. Piles suspected of damage (including timber, H, and 

solid concrete piles) could be tested to confirm integrity and/or determine extent and 

location of damage using high strain dynamic pile testing, or for concrete piles, low 

strain integrity testing methods. These methods are discussed in Chapter 10. 

Alternatively, the pile could be replaced or repaired, if possible.

Check for water leakage and soil inflow into closed end pipe piles before placing 

concrete. The concrete mix should have a high slump and small aggregate. A pipe 

pile can be easily checked for damage and sweep by lowering a light source inside 

the pile. A mirror can also be used to reflect sunlight inside a pipe pile for internal 

inspection.
496

The driving sequence of piles in a pier or bent can be important. The driving 

sequence can affect the way piles drive as well as the influence the new 

construction has on adjacent structures. This is especially true for displacement 

piles. For non-displacement piles, the driving sequence is generally not as critical.

The driving sequence of displacement pile groups should be from the center of the 

group outward or from one side to the other side. The preferred driving sequence of 

the displacement pile group shown in Figure 18-18 would be (a) by the pile number 

shown, (sequence 1), (b) by driving each row starting in the center and working 

outward (sequence 2), or (c) by driving each row starting on one side of the group 

and working to the other side (sequence 3).

Figure 18-18 Driving sequence of displacement pile groups (after Passe 1994).

Pile groups should not be driven from the outside to the center (the reverse of 

sequences 1 or 2). If groups are driven in that order, displaced soils becomes 

trapped and compacted in the center of the pile group. This can cause problems 
497

with driving the piles in the center of the group and prevent those piles from meeting 

minimum pile penetration depths for scour.

When driving close to an existing structure or utility, it is generally preferable to drive 

the piles nearest the existing structure or utility first and work away. For example, if 

a structure was located on the right side of the pile group shown in Figure 18-18, the 

piles should be driven by sequence 3. This reduces the amount of soil displaced 

toward the existing structure. The displacement of soil toward an existing structure 

has caused problems before. It can be especially critical next to a bascule bridge 

where, very small movements can prevent the locking mechanism from locking.

On some projects, vibration measurements may be required to ascertain if pile 

driving vibrations are within acceptable and/or specified maximum levels. Woods 

(1997) noted that vibration damage is relatively uncommon at a distance of one pile 

length away from driving. However, damage from vibration induced settlement of 

loose, clean sands can be a problem up to 1300 feet away from driving. Prior to the 

start of construction, a survey of structures and utilities within 400 feet of pile driving 

activities is often performed to document their existing condition. Some 

specifications tie the preconstruction survey limits to the hammer energy and require 

a preconstruction survey of all structures located within a distance in feet of 0.25 

times the square root of the hammer energy in foot-pounds. The preconstruction 

survey generally consists of photographing or videotaping existing damage, as well 

as affixing crack gages to existing cracks in some cases.

Woods (1997) noted that damage to freshly placed concrete from pile driving 

vibrations may not be a risk but further research on the setting and curing of 

concrete may be warranted.

A cold hammer should not be used when restriking piles after a setup period. 

Twenty hammer blows are usually sufficient to warm up most hammers. Also be 

sure to record the restrike penetration resistance for each 1 inch interval during the 

restrike.

A summary of common pile installation problems and possible solutions is presented 

in Table 18-7. 
498

Table 18-7 Common Pile Installation Problems and Possible Solutions

Common Problems Possible Solutions

Piles encountering 

refusal blow count above 

minimum pile 

penetration 

requirements.

Have wave equation analysis performed and check 

that pile has sufficient drivability and that the driving 

system is matched to the pile. If the pile and driving 

system are suitably matched, check driving system 

operation for compliance with manufacturer's 

guidelines. If no obvious problems are found, 

dynamic measurements should be made to 

determine if the problem is driving system or soil 

behavior related. Driving system problems could 

include pre-ignition, preadmission, low hammer 

efficiency, or soft cushion. Soil problems could 

include greater soil strength than anticipated, 

temporarily increased soil resistance with later 

relaxation (requires restrike to check), large soil 

quakes, or high soil damping. 

Piles driving significantly 

deeper than estimated 

pile penetration depths.

Soil resistance at the time of driving probably is 

lower than anticipated or driving system 

performance is better than anticipated. Have wave 

equation analysis performed to assess nominal 

resistance based on the blow count at the time of 

driving. Perform restrike tests after an appropriate 

waiting period to evaluate soil strength changes 

with time. If the nominal resistance based on 

restrike blow count is still low, check drive system 

performance and restrike pile with dynamic 

measurements. If drive system performance is as 

assumed and restrike nominal resistance low, the 

soil conditions are weaker than anticipated. 

Foundation piles will probably need to be driven 

deeper than originally estimated or additional piles 

will be required to support the load. Contact the 

structural engineer/designer for recommended 

change.
499

Table 18-7 Common Pile Installation Problems and Possible Solutions (Continued)

Common Problems Possible Solutions

Abrupt change or 

decrease in blow count 

for bearing piles.

If borings do not indicate weathered profile above 

bedrock/bearing layer then pile toe damage is likely. 

Have wave equation analysis performed and 

evaluate pile toe stress. If calculated toe stress is 

high and blow counts are low, a reduced hammer 

energy (stroke) and higher blow count could be 

used to achieve nominal resistance with a lower toe 

stress. If calculated toe stress is high at high blow 

counts, a different hammer or pile section may be 

required. For piles that allow internal inspection, 

reflect light to the pile toe and tape the length inside 

the pile for indications of toe damage. For piles that 

cannot be internally inspected, dynamic 

measurements could be made to evaluate problem 

or pile extraction could be considered for 

confirmation of a damage problem.

Blow count significantly 

lower than expected 

during driving.

Review soil borings. If soil borings do not indicate 

soft layers, pile may be damaged below grade. 

Have wave equation analysis performed and 

investigate both tensile stresses along pile and 

compressive stresses at toe. If calculated stresses 

are within allowable limits, investigate possibility of 

obstructions / uneven toe contact on hard layer or 

other reasons for pile toe damage. If pile was 

spliced, re-evaluate splice detail and field splicing 

procedures for possible splice failure.

Vertical (heave) or 

lateral movement of 

previously installed piles 

when driving new piles.

Pile movements likely due to soil displacement from 

adjacent pile driving. Contact geotechnical 

engineer for recommended action. Possible 

solutions include redriving of installed piles, change 

in sequence of pile installation, or predrilling of pile 

locations to reduce ground movements. Lateral pile 

movements could also result from adjacent slope 

failure in applicable conditions.
500

Table 18-7 Common Pile Installation Problems and Possible Solutions (Continued)

Common Problems Possible Solutions

Piles driving out of 

alignment tolerance.

Piles may be moving out of alignment tolerance due 

to hammer-pile alignment control or due to soil 

conditions. If due to poor hammer-pile alignment 

control, a pile gate, template or fixed lead system 

may improve the ability to maintain alignment 

tolerance. Soil conditions such as near surface 

obstructions (see subsequent section) or steeply 

sloping bedrock having minimal overburden 

material (pile point detail is important) may prevent 

tolerances from being met even with good 

alignment control. In these cases, survey the asbuilt condition and contact the structural engineer 

for recommended action.

Piles driving out of 

location tolerance.

Piles may be moving out of location tolerance due 

to hammer-pile alignment control or due to soil 

conditions. If due to poor hammer-pile alignment 

control, a pile gate, template or fixed lead system 

may improve the ability to maintain location 

tolerance. Soil conditions such as near surface 

obstructions (see subsequent section) or steeply 

sloping bedrock having minimal overburden 

material (pile point detail is important) may prevent 

tolerances from being met even with good 

alignment control. In these cases, survey the asbuilt condition and contact the structural engineer 

for recommended action. 

Piles encountering 

shallow obstructions.

If obstructions are within 10 feet of working grade, 

obstruction excavation and removal is probably 

feasible. If obstructions are at deeper depth, are 

below the water table, or the soil is contaminated, 

excavation may not be feasible. Spudding or 

predrilling of pile locations may provide a solution 

with method selection based on the type of 

obstructions and soil conditions.
501

Table 18-7 Common Pile Installation Problems and Possible Solutions (Continued)

Common Problems Possible Solutions

Piles encountering 

obstructions at depth.

If deep obstructions are encountered that prevent 

reaching the desired pile penetration depth, contact 

the structural engineer/designer for remedial 

design. Nominal resistance of piles hitting 

obstructions should be reduced based upon pile 

damage potential and soil matrix support 

characteristics. Additional foundation piles may be 

necessary. 

Concrete piles develop 

complete horizontal 

cracks in easy driving.

Have wave equation analysis performed and check 

tension stresses along pile (extrema tables) for the 

observed blow counts. If the calculated tension 

stresses are high, add cushioning or reduce stroke. 

If calculated tension stresses are low, check 

hammer performance and/or perform dynamic 

measurements.

Concrete piles develop 

complete horizontal 

cracks in hard driving.

Have wave equation analysis performed and check 

tension stresses along pile (extrema table). If the 

calculated tension stresses are high, consider a 

hammer with a heavier ram. If the calculated 

tension stresses are low, perform dynamic 

measurements and evaluate soil quakes which are 

probably higher than anticipated.

Concrete piles develop 

partial horizontal cracks 

in easy driving.

Check hammer-pile alignment since bending may 

be causing the problem. If the alignment appears to 

be normal, tension and bending combined may be 

too high. The possible solution is as above with 

complete cracks.

Concrete pile spalling or 

slabbing near pile head.

Have wave equation analysis performed. 

Determine the pile head stress at the observed blow 

count and compare predicted stress with material 

stress limits. If the calculated stress is high, 

increase the pile cushioning. If the calculated 

stress is low, investigate pile quality, hammer 

performance, and hammer-pile alignment.
502

Table 18-7 Common Pile Installation Problems and Possible Solutions (Continued)

Common Problems Possible Solutions

Steel pile head deforms. Check helmet size/shape, steel yield strength, and 

evenness of the pile head. If all seem acceptable, 

have wave equation analysis performed and 

determine the pile head stress. If the calculated 

stress is high and blow counts are low, use reduced 

hammer energy (stroke) and higher blow count to 

achieve nominal resistance. If the calculated stress 

is high at high blow counts, a different hammer or 

pile type may be required. Nominal resistance 

determination should not be made using blow 

counts obtained when driving with a deformed pile 

head.

Timber pile head 

mushrooms. 

Check helmet size/shape, the evenness of the pile 

head, and banding of the timber pile head. If all 

seem acceptable, have wave equation analysis 

performed and determine the pile head stress. If 

the calculated stress is high and blow counts are 

low, use reduced hammer energy (stroke) and 

higher blow count to achieve nominal resistance. 

Nominal resistance determination should not be 

made using blow counts obtained when driving with 

a mushroomed pile head.

18.10 DRIVING RECORDS AND REPORTS

Pile driving records vary depending upon the organization performing the monitoring 

or inspection service. A typical pile driving record is presented in Figure 18-19. The 

following is a list of items that appear on most pile driving records:

1. Project identification number.

2. Project name and location.

3. Structure identification number.

4. Date and time of driving (start, stop, and interruptions).
503

5. Name of the contractor. 

6. Hammer make, model, ram weight, energy rating. The actual stroke and 

operating speed should also be recorded whenever it is changed.

7. Hammer cushion description, size and thickness, and helmet weight.

8. Pile cushion description, size and thickness, depth where changed.

9. Pile location, type, size and length.

10. Pile number or designation matching pile layout plans.

11. Pile ground surface, cut off, and final penetration elevations and embedded 

length.

12. Penetration resistance data in blows per foot with the final 1 foot normally 

recorded in blows per inch.

13. Graphical presentation of driving data (optional).

14. Cut off length, length in ground and order length.

15. Comments or unusual observations, including reasons for all interruptions.

16. Signature and title of the inspector.

The importance of maintaining detailed pile driving records cannot be 

overemphasized. The driving records form a basis for payment and for making 

engineering decisions regarding the adequacy of the foundation to support the 

design loads. Great importance is given to driving records in litigations involving 

claims. Sloppy, inaccurate, or incomplete records encourage claims and result in 

higher cost foundations. The better the pile driving is documented, the lower the 

foundation cost will probably be and the more likely it will be completed on schedule. 

In addition to the driving records, the inspector should be required to prepare a daily 

inspection report. The daily inspection report should include information on 

equipment working at the site, description of construction work accomplished, and 

the progress of work. Figure 18-20 shows an example of a daily inspection report.
504

Figure 18-19 Pile driving log.
505

Figure 18-20 Daily inspection report.
506

18.11 SAFETY

Pile driving involves the use of heavy equipment and heavy loads. The pile driving 

inspector should be cognizant of these activities and position his or herself 

accordingly. One of the more dangerous operations during pile driving is the lifting 

of the pile and the positioning of it under the hammer for driving. The inspector 

should remember that a 100 feet pile can fall 100 feet from the pile location during 

positioning. It is better to have a planned escape route prior to pile positioning rather 

than attempt to quickly determine one should difficulties arise during the pile lifting 

and positioning process.

The area beneath a suspended load should be avoided. If the hoisting device fails 

or slips, serious injury could occur. The inspector should also avoid the area behind 

the crane and be cognizant whenever the crane is moving or swinging.

The inspector should select a position to maintain the pile driving record that is a 

sufficient distance away from the pile location during driving. The area immediately 

in front of the pile should be avoided. Heavy pieces sometimes fall from a pile 

hammer or helmet during operation that could cause serious injury if the inspector 

were positioned under or near the hammer. All pile types can be also damaged 

during driving. Concrete and timber piles can break suddenly, and long steel piles 

can buckle. A sudden pile break or buckling can make the area around the pile 

location quite dangerous due to the broken or buckled section. A sudden loss of 

resistance beneath an operating pile hammer can also overload the hammer line 

causing it to break and the hammer to fall. Damage to the head of a concrete pile 

during driving can also result in heavy concrete pieces falling due to hammer-pile 

alignment problems or due to pile cushion deterioration. Standing beneath the 

hammer and monitoring the final pile penetration resistance by placing marks on the 

pile every 10 or 20 hammer blows should be avoided for the above reasons. The 

final blow count can be determined instead by marking the pile prior to driving with 1 

inch marks over the anticipated final penetration depth.

Safety devices such as a hard hat, ear protection, steel toed work boots, eye 

protection, safety vest, fall protection harness, and life jacket should be worn as job 

conditions dictate. Additional site specific, state, or federal safety rules may also 

apply and these should be reviewed.
507

REFERENCES

Deep Foundations Institute (DFI). (1995). A Pile Inspector's Guide to Hammers, 

Second Edition, Deep Foundations Institute, Englewood Cliffs, NJ, 71 p. 

Deep Foundations Institute (DFI). (1997). Inspectors Manual for Driven Pile 

Foundations, Second Edition, Deep Foundations Institute, Englewood Cliffs, 

NJ, 69 p. 

Passe, Paul D. (1994). Pile Driving Inspector's Manual. State of Florida Department 

of Transportation.

Rausche, F., Likins, G.E., Goble, G.G., Hussein, M. (1986). The Performance of Pile 

Driving Systems; Inspector's Manual. U.S. Department of Transportation, 

Federal Highway Administration, Washington, D.C., 92 p.

Williams Earth Sciences (1995). Inspector's Qualification Program for Pile Driving 

Inspection Manual. State of Florida Department of Transportation.

Woods, R.D. (1997). Dynamic Effects of Pile Installations on Adjacent Structures. 

NCHRP Synthesis 253, National Cooperative Highway Research Program, 

Transportation Research Board, Washington, D.C., 85 p.
508
509

Appendix A

LIST OF FHWA/ NHI RESOURCES RELEVANT TO DEEP FOUNDATIONS

Abu-Hejleh, N., DiMaggio, J.A., Kramer, W.M., Anderson, S., and Nichols, S. (2010). 

Implementation of LRFD Geotechnical Design for Bridge Foundations: 

Reference Manual. FHWA-NHI-10-039. National Highway Institute, Federal 

Highway Administration, Washington, D.C.

Abu-Hejleh, N., Kramer, W.M., Mohamed, K., Long, J.H., and Zaheer, M.A. (2013). 

Implementation of AASHTO LRFD Design Specifications for Driven Piles, 

FHWA-RC-13-001. U.S. Dept. of Transportation, Federal Highway 

Administration, Washington, D.C., 71 p.

Arneson, L.A., Zevenbergen, L.W., Lagasse, P.F., and Clopper, P.E. (2012). 

Evaluating Scour at Bridges, Fifth Edition, FHWA-HIF-12-003, Hydraulic 

Engineering Circular (HEC) No. 18. U.S. Dept. of Transportation, Federal 

Highway Administration, 340 p.

Azizinamini, A., Power, E.H., Myers, G.F., and Ozyildirim, H.C. (2014). Bridges for 

Service Life Beyond 100 Years, Innovative Systems Subsystems, and 

Components, S2-R19A-RW-1 Strategic Highway Research Program 2 

(SHRP), Transportation Research Board, Washington, D.C., 248 p.

Briaud J-L. and Miran, J. (1992). The Cone Penetration Test, FHWA-SA-91-043. 

U.S. Department of Transportation, Federal Highway Administration, Office of 

Technology Applications, Washington, D.C., 161 p.

Brown, D.A., and Thompson, W.R. (2011). Developing Production Pile Criteria from 

Test Pile Data. National Cooperative Highway Research Program (NCHRP) 

Synthesis 418, Washington, D.C., 54 p. 

Brown, D. A., Turner, J.P. and Castelli R.J. (2010). Drilled Shafts: Construction 

Procedures and LRFD Design Methods, FHWA-NHI-10-016, Geotechnical 

Engineering Circular (GEC) No. 10. U.S. Dept. of Transportation, Federal 

Highway Administration, 970 p.
510

Brown, D.A., Dapp, S.D., Thompson, W.R., and Lazarte, C.A. (2007). Design and 

Construction of Continuous Flight Auger (CFA) Piles. FHWA-HIF-07-03, 

Geotechnical Engineering Circular (GEC) No.08. U.S. Dept. of 

Transportation, Federal Highway Administration, 289 p.

Cheney, R.S. and Chassie, R.G. (2000). Soils and Foundations Workshop 

Reference Manual. FHWA HI-00-045, U.S. Department of Transportation, 

National Highway Institute, Federal Highway Administration, Washington, 

D.C., 358 p.

Elias, V., Welsh, J.P., Warren, J., Lukas, R.G., Collin J.G., and Berg, R.R. (2006). 

Ground Improvement Methods Volumes I and II, FHWA-NHI-06-019 and 

FHWA NHI-06-020. National Highway Institute, Federal Highway 

Administration, U.S. Department of Transportation, Washington D.C.

Federal Highway Administration (FHWA). (1996). Geotechnical Engineering 

Notebook DT-15. Differing Site Conditions. U.S. Dept. of Transportation, 

Federal Highway Administration, Washington, D.C., 36 p.

Federal Highway Administration (FHWA). (2002a). Life-Cycle Cost Analysis Primer, 

IF-02-047. Federal Highway Administration, U.S. Department of 

Transportation. Washington, D.C., 24 p.

Geotechnical Guideline 13 (1985). Geotechnical Engineering Notebook, U.S. 

Department of Transportation. Federal Highway Administration, Washington, 

D.C., 37 p.

Ghosn, M., Moses, F., and Wang, J. (2003). Design of Highway Bridges for Extreme 

Events, NCHRP Report 489. National Cooperative Highway Research 

Program. Transportation Research Board, Washington, D.C., 174 p.

Goble, G.G. and Rausche, F. (1976). Wave Equation Analysis of Pile Driving – 

WEAP Program, FHWA IP-76-14.3., U.S. Department of Transportation, 

Federal Highway Administration, Office of Research and Development, 

Washington, D.C., Volumes I-IV.

Goble, G.G. and Rausche, F. (1986). Wave Equation Analysis of Pile Driving - 

WEAP86 Program. U.S. Department of Transportation, Federal Highway 

Administration, Implementation Division, McLean, VA, Volumes I-IV.
511

Hawk, H. (2003). Bridge Life-Cycle Cost Analysis, NCHRP483. Transportation 

Research Board of the National Academies, Washington, D.C., 96 p.

Kavazanjian, E., Wan, J-N. J., Martin, G.R., Shamsabadi, A., Lam, I., Dickenson, 

S.E., and Hung, C.J. (2011). LRFD Seismic Analysis and Design of 

Transportation Geotechnical Features and Structural Foundations, FHWANHI-11-032, Geotechnical Engineering Circular (GEC) No. 3. U.S. Dept. of 

Transportation, Federal Highway Administration, Washington, D.C., 592 p.

Kimmerling, R.E. (2002). Shallow Foundations, FHWA-IF-02-054, Geotechnical 

Engineering Circular (GEC) No. 6. U.S. Dept. of Transportation, Federal 

Highway Administration, Washington, D.C., 310 p.

Kyfor, Z.G., Schnore, A.R., Carlo, T.A. and Bailey, P.F. (1992). Static Testing of 

Deep Foundations, FHWA-SA-91-042. U.S. Department of Transportation, 

Federal Highway Administration, Office of Technology Applications, 

Washington, D.C., 174 p.

Lam, I.P. and Martin, G.R. (1986). Seismic Design of Highway Bridge Foundations. 

Volume II - Design Procedures and Guidelines, FHWA-RD-86-102. U.S. 

Department of Transportation, Federal Highway Administration, Office of 

Engineering and Highway Operations, McLean, VA, 167p.

Long, J., and Anderson, A. (2014). Improved of Driven Pile Installation and Design in 

Illinois: Phase 2, FHWA-ICT-14-019. Illinois Department of Transportation, 

Bureau of Material and Physical Research, Springfield, IL, 84 p.

Marsh, M.L., Buckle, I.G., and Kavazanjian Jr, E. (2014). LRFD Seismic Analysis 

and Design of Bridges, FHWA-NHI-15-004. National Highway Institute, U.S. 

Dept. of Transportation, Federal Highway Administration, Washington, D.C., 

608 p.

Mayne, P.W., Christopher, B., Berg, R., and DeJong, J. (2001). Manual on 

Subsurface Investigations, FHWA NHI-01-031, U.S. Dept. of Transportation, 

National Highway Institute, Federal Highway Administration, Washington, 

D.C., 301 p.

Munfakh, G., Arman, A., Collin, J.G., Hung, J.C.-J., and Brouillette, R.P. (2001). 

Shallow Foundations Reference Manual, FHWA-NHI-01-023. National 

Highway Institute, Federal Highway Administration, Washington, D.C., 222 p.
512

Rausche, F., Likins, G.E., Goble, G.G., Hussien, M. (1986). The Performance of Pile 

Driving Systems; Inspector's Manual. U.S. Department of Transportation, 

Federal Highway Administration, Washington, D.C., 92 p.

Rausche, F., Thendean, G., Abou-matar, H., Likins, G.E. and Goble, G.G. (1996). 

Determination of Pile Drivability and Capacity from Penetration Tests, 

DTFH61-91-C-00047, Final Report. U.S. Department of Transportation, 

Federal Highway Administration, McLean, VA, 432 p.

Reese, L.C. (1984). Handbook on Design of Piles and Drilled Shafts Under Lateral 

Load. Report No. FHWA-IP-84-11, U.S. Department of Transportation, 

Federal Highway Administration, Office of Implementation, McLean, VA, 386 

p.

Reese, L.C. (1986). Behavior of Piles and Pile Groups Under Lateral Load, FHWARD-85-106. U.S. Department of Transportation, Federal Highway 

Administration, Office of Engineering and Highway Operations Research and 

Development, Washington, D.C., 311 p.

Rixner, J.J., Kraemer, S.R. and Smith, A.D. (1986). Prefabricated Vertical Drains 

Volume I, Engineering Guidelines, FHWA-RD-86-168. U.S. Department of 

Transportation, Federal Highway Administration, Office of Engineering and 

Highway Operations Research and Development, McLean, VA, 117 p.

Sabatini, P. J., Elias, V., Schmertmann, G. R., and Bonaparte, R. (1997). Earth 

Retaining Systems FHWA-SA-96-038, Geotechnical Engineering Circular 

(GEC) No. 2. U.S. Department of Transportation, Federal Highway 

Administration, Washington, D.C.

Sabatini, P.J., Tanyu, B., Armour., P., Groneck, P., and Keeley, J. (2005). Micropile 

Design and Construction, FHWA-NHI-05-039. National Highway Institute, 

U.S. Dept. of Transportation, Federal Highway Administration, Washington, 

D.C., 436 p.

Samtani, N.C. and Nowatzki, E.A. (2006). Soils and Foundations: Reference 

Manual, Vol. 1, FHWA-NHI-06-088, U.S. Dept. of Transportation, National 

Highway Institute, Federal Highway Administration, Washington, D.C., 462 p.
513

Samtani, N.C., Nowatzki, E.A., and Mertz, D.R. (2010). Selection of Spread Footings 

on Soils to Support Highway Bridge Structures, FHWA-RC TD-10-001. U.S. 

Department of Transportation, Federal Highway Administration, Washington, 

D.C., 98 p.

Schmertmann, J.H. (1978). Guidelines For Cone Penetration Test, Performance, 

and Design, FHWA-TS-78-209. U.S. Department of Transportation, Federal 

Highway Administration, Washington, D.C., 145 p.

Tanyu B.F., Sabatini, P. J., and Berg, R.R. (2008). Earth Retaining Structures, 

FHWA-NHI-07-07. U.S. Department of Transportation, Federal Highway

Administration, Washington, D.C., 792 p.

Wightman, W., Jalinoos, F., Sirles., and Hanna, K. (2003), Applications of 

Geophysical Methods to Related Highway Problems. U.S. Dept. of 

Transportation, Federal Highway Administration, 716 p.

Wilson, K.E., Kimmerling, R.E., Goble, G.C., Sabatini, P.J., Zang, S.D., Zhou, J.Y. 

Amrhein, W.A., Bouscher, J.W. and Danaovich, L.J. (2006). LRFD for 

Highway Bridge Substructures and Earth Retaining Structures Reference 

Manual, FHWA NHI-05-094. U.S. Dept. of Transportation, Federal Highway 

Administration, Washington, D.C., 1730 p.
514
515

Appendix B

LIST OF ASTM AND AASHTO PILE DESIGN AND TESTING SPECIFICATIONS

American Association of State Highway and Transportation Offices (AASHTO). 

Standard Method of Test for High Strain Dynamic Testing of Piles, AASHTO 

Designation T-298-33.

American Association of State Highway and Transportation Officials (AASHTO). 

(1978). Manual on Foundation Investigations Second Edition.. AASHTO 

Highway Subcommittee on Bridges and Structures, Washington, D.C., 196 p.

American Association of State Highway and Transportation Officials (AASHTO). 

(2001). Standard Recommended Practice for Assessment of Corrosion of 

Steel Piling for Non-Marine Applications. AASHTO Standard Specifications 

for Transportation Materials and Methods of Sampling and Testing, Part 1B: 

Specifications, 24th Edition,. 13 p.

American Association of State Highway and Transportation Officials (AASHTO). 

(2011) Guide Specifications for LRFD Seismic Bridge Design, 2nd Edition, 

with 2012, 2014, and 2015 Interim Revisions. American Association of State 

Highway and Transportation Officials, Washington, D.C., 331 p.

American Association of State Highway and Transportation Officials (AASHTO). 

(2014). AASHTO LRFD Bridge Design Specifications, US Customary Units, 

Seventh Edition, with 2015 Interim Revisions. American Association of State 

Highway and Transportation Officials, Washington, D.C., 1960 p.

ASTM A27-13. (2014). Standard Specification for Steel Castings, Carbon, for 

General Application. Book of ASTM Standards, Vol. 1.02, ASTM International, 

West Conshohocken, PA, 4 p. 

ASTM A572-15. (2015). Standard Specification for High-Strength Low-Alloy 

Columbium-Vanadium Structural Steel. Book of ASTM Standards, Vol. 1.04, 

ASTM International, West Conshohocken, PA, 4 p.
516

ASTM D1143-07. (2014). Standard Test Methods for Deep Foundations Under 

Static Axial Compressive Load. Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 15 p.

ASTM D1452-09. (2014). Standard Practice for Soil Exploration and Sampling by 

Auger Borings. Annual Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 6 p.

ASTM D1586-11. (2014). Standard Test Method for Standard Penetration Test 

(SPT) and Split-Barrel Sampling of Soils. Annual Book of ASTM Standards, 

Vol. 4.08, ASTM International, West Conshohocken, PA, 9 p.

ASTM D1587-12. (2014). Standard Practice for Thin-Walled Tube Sampling of Soils 

for Geotechnical Purposes. Annual Book of ASTM Standards, Vol. 4.08, 

ASTM International, West Conshohocken, PA, 4 p.

ASTM D2113-14. (2014). Standard Practice for Rock Core Drilling and Sampling of 

Rock for Site Investigation. Annual Book of ASTM Standards, Vol. 4.08, 

ASTM International, West Conshohocken, PA, 20 p.

ASTM D2573-08. (2012). Standard Test Method for Field Vane Shear Test in 

Cohesive Soil. Annual Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 8 p.

ASTM D3689-07. (2014). Standard Test Methods for Deep Foundations Under 

Static Axial Tensile Load. Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 13 p.

ASTM D4633-10. (2014). Standard Test Method for Energy Measurement for 

Dynamic Penetrometer. Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 7 p.

ASTM D4719-07. (2014). Standard Test Methods for Prebored Pressuremeter 

Testing in Soils Annual. Book of ASTM Standards, Vol. 4.08, ASTM 

International, West Conshohocken, PA, 10 p.

ASTM D4945-12. (2014). Standard Test Method for High-Strain Dynamic Testing of 

Piles. Book of ASTM Standards, Vol. 4.08, ASTM International, West 

Conshohocken, PA, 9 p.
517

ASTM D4971-02. (2014). Standard Test Method for Determining the In Situ Modulus 

of Deformation of Rock Using the Diametrically Loaded 76-mm (3-in.) 

Borehole Jack. Book of ASTM Standards, Vol. 4.08, ASTM International, 

West Conshohocken, PA, 7 p.

ASTM D5778-12. (2014). Standard Test Method for Electronic Friction Cone and 

Piezocone Penetration Testing of Soils. Book of ASTM Standards, Vol. 4.08, 

ASTM International, West Conshohocken, PA, 20 p.

ASTM D5882-07. (2013). Standard Test Method for Low-Strain Dynamic Testing of 

Piles Annual Book of ASTM Standards, Vol. 4.09, ASTM International, West 

Conshohocken, PA, 6 p.

ASTM D6032-08. (2014). Standard Test Method for Determining Rock Quality 

Designation (RQD) of Rock Core. Annual Book of ASTM Standards, Vol. 

4.09, ASTM International, West Conshohocken, PA, 5 p.

ASTM D6635-07. (2014). Standard Test Method for Performing the Flat Dilatometer. 

Book of ASTM Standards, Vol. 4.09, ASTM International, West

Conshohocken, PA, 16 p.

ASTM D7012-14. (2014). Standard Tests Method for Compressive Strength and 

Elastic Moduli of Intact Rock Core Specimens under Varying States of Stress 

and Temperatures. Annual Book of ASTM Standards, Vol. 4.09, ASTM 

International, West Conshohocken, PA, 9 p.

ASTM D7383-10 (2010). Standard Test Methods for Axial Compressive Force Pulse 

(Rapid) Testing of Deep Foundations. Annual Book of ASTM Standards, Vol. 

4.08, ASTM International, West Conshohocken, PA, 9 p.

ASTM Vol 4.08. (2014). Soil and Rock I, Vol. 4.08, ASTM International, West 

Conshohocken, PA, 1826 p.

ASTM Vol 4.09. (2014). Soil and Rock II, Vol. 4.09, ASTM International, West 

Conshohocken, PA, 1754 p.
518
519

Appendix C

PILE HAMMER INFORMATION

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

1 DELMAG D 5 OED 10.51 1.10 9.55

2 DELMAG D 8-22 OED 20.10 1.76 11.42

3 DELMAG D 12 OED 22.61 2.75 8.22

4 DELMAG D 15 OED 27.09 3.30 8.21

5 DELMAG D 16-32 OED 40.20 3.52 11.42

6 DELMAG D 22 OED 40.61 4.91 8.27

7 DELMAG D 22-02 OED 48.50 4.85 10.00

8 DELMAG D 22-13 OED 48.50 4.85 10.00

9 DELMAG D 22-23 OED 51.22 4.85 10.56

10 DELMAG D 25-32 OED 66.34 5.51 12.04

11 DELMAG D 30 OED 59.73 6.60 9.05

12 DELMAG D 30-02 OED 66.20 6.60 10.03

13 DELMAG D 30-13 OED 66.20 6.60 10.03

14 DELMAG D 30-23 OED 73.79 6.60 11.18

15 DELMAG D 30-32 OED 75.44 6.60 11.43

16 DELMAG D 36 OED 83.82 7.93 10.57

17 DELMAG D 36-02 OED 83.82 7.93 10.57

18 DELMAG D 36-13 OED 83.82 7.93 10.57

19 DELMAG D 36-23 OED 88.50 7.93 11.16

20 DELMAG D 36-32 OED 90.56 7.93 11.42

21 DELMAG D 44 OED 90.16 9.50 9.49

22 DELMAG D 46 OED 107.08 10.14 10.56

23 DELMAG D 46-02 OED 107.08 10.14 10.56

24 DELMAG D 46-13 OED 96.53 10.14 9.52

25 DELMAG D 46-23 OED 107.08 10.14 10.56

26 DELMAG D 46-32 OED 122.19 10.14 12.05

27 DELMAG D 55 OED 125.00 11.86 10.54

28 DELMAG D 62-02 OED 152.45 13.66 11.16

29 DELMAG D 62-12 OED 152.45 13.66 11.16

30 DELMAG D 62-22 OED 164.60 13.66 12.05

31 DELMAG D 80-12 OED 186.24 17.62 10.57

32 DELMAG D 80-23 OED 212.50 17.62 12.06

33 DELMAG D100-13 OED 265.68 22.07 12.04

35 DELMAG D 19-52 OED 43.20 4.00 10.80

36 DELMAG D 6-32 OED 13.52 1.32 10.23
520

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

37 DELMAG D 12-32 OED 31.33 2.82 11.11

38 DELMAG D 12-42 OED 33.30 2.82 11.81

39 DELMAG D 14-42 OED 34.50 3.09 11.18

40 DELMAG D 19-32 OED 42.44 4.00 10.61

41 DELMAG D 19-42 OED 43.24 4.00 10.81

42 DELMAG D200-42 OED 492.04 44.09 11.16

43 DELMAG D120-42 OED 301.79 26.45 11.41

44 DELMAG D150-42 OED 377.33 33.07 11.41

45 DELMAG D125-42 OED 313.63 27.56 11.38

46 DELMAG D 21-42 OED 55.75 4.63 12.04

47 DELMAG D 5-42 OED 10.56 1.10 9.60

48 DELMAG D160-32 OED 393.45 35.27 11.16

49 DELMAG D260-32 OED 639.36 57.32 11.16

50 FEC 1200 OED 22.50 2.75 8.18

51 FEC 1500 OED 27.09 3.30 8.21

52 FEC 2500 OED 50.00 5.50 9.09

53 FEC 2800 OED 55.99 6.16 9.09

54 FEC 3000 OED 63.03 6.60 9.55

55 FEC 3400 OED 73.01 7.48 9.76

56 FEC D-18 OED 39.70 3.97 10.00

61 MITSUBIS M 14 OED 25.25 2.97 8.50

62 MITSUBIS MH 15 OED 28.14 3.31 8.50

63 MITSUBIS M 23 OED 43.01 5.06 8.50

64 MITSUBIS MH 25 OED 46.84 5.51 8.50

65 MITSUBIS M 33 OED 61.71 7.26 8.50

66 MITSUBIS MH 35 OED 65.62 7.72 8.50

67 MITSUBIS M 43 OED 80.41 9.46 8.50

68 MITSUBIS MH 45 OED 85.43 10.05 8.50

70 MITSUBIS MH 72B OED 135.15 15.90 8.50

71 MITSUBIS MH 80B OED 149.60 17.60 8.50

81 LINKBELT LB 180 CED 8.10 1.73 4.68

82 LINKBELT LB 312 CED 15.02 3.86 3.89

83 LINKBELT LB 440 CED 18.20 4.00 4.55

84 LINKBELT LB 520 CED 26.31 5.07 5.19

85 LINKBELT LB 660 CED 51.63 7.57 6.82

90 HITACHI HNC65 ECH 56.42 14.33 3.94 

91 HITACHI HNC80 ECH 69.43 17.64 3.94

92 HITACHI HNC100 ECH 86.79 22.05 3.94

93 HITACHI HNC125 ECH 108.49 27.56 3.94

101 KOBE K 13 OED 25.43 2.87 8.86

103 KOBE K22-Est OED 45.35 4.85 9.35

104 KOBE K 25 OED 51.52 5.51 9.35

107 KOBE K 35 OED 72.18 7.72 9.35
521

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

110 KOBE K 45 OED 92.75 9.92 9.35

112 KOBE KB 60 OED 130.18 13.23 9.84

113 KOBE KB 80 OED 173.58 17.64 9.84

120 ICE 180 CED 8.13 1.73 4.70

121 ICE 422 CED 23.12 4.00 5.78

122 ICE 440 CED 18.56 4.00 4.64

123 ICE 520 CED 30.37 5.07 5.99

124 ICE 640 CED 40.62 6.00 6.77

125 ICE 660 CED 51.63 7.57 6.82

126 ICE 1070 CED 72.60 10.00 7.26

127 ICE 30-S OED 22.50 3.00 7.50

128 ICE 40-S OED 40.00 4.00 10.00

129 ICE 42-S OED 42.00 4.09 10.27

130 ICE 60-S OED 59.99 7.00 8.57

131 ICE 70-S OED 70.00 7.00 10.00

132 ICE 80-S OED 80.00 8.00 10.00

133 ICE 90-S OED 90.00 9.00 10.00

134 ICE 100-S OED 100.00 10.00 10.00

135 ICE 120-S OED 120.00 12.00 10.00

136 ICE 200-S OED 100.00 20.00 5.00

137 ICE 205-S OED 170.00 20.00 8.50

139 ICE 32-S OED 26.01 3.00 8.67

140 ICE 120S-15 OED 132.45 15.00 8.83

142 MKT DE-20C OED 20.00 2.00 10.00

143 MKT DE-30C OED 28.00 2.80 10.00

144 MKT DE-33C OED 33.00 3.30 10.00

145 MKT DE333020 OED 40.00 4.00 10.00

146 MKT DE 10 OED 8.80 1.10 8.00

147 MKT DE 20 OED 16.00 2.00 8.00

148 MKT DE 30 OED 22.40 2.80 8.00

149 MKT DA35B SA OED 23.80 2.80 8.50

150 MKT DE 30B OED 23.80 2.80 8.50

151 MKT DA 35B CED 21.00 2.80 7.50

152 MKT DA 45 CED 30.72 4.00 7.68

153 MKT DE 40 OED 32.00 4.00 8.00

154 MKT DE 35 OED 35.00 3.50 10.00

155 MKT DE 42 OED 42.00 4.20 10.00

157 MKT DE 50C OED 50.00 5.00 10.00

158 MKT DE 70C OED 70.00 7.00 10.00

159 MKT DE 50B OED 42.50 5.00 8.50

160 MKT DA55B SA OED 40.00 5.00 8.00

161 MKT DA 55B CED 38.20 5.00 7.64

162 MKT DE 70B OED 59.50 7.00 8.50
522

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

163 MKT DE-50B OED 50.00 5.00 10.00

164 MKT DE-70B OED 70.00 7.00 10.00

165 MKT DE-110C OED 110.00 11.00 10.00

166 MKT DE-150C OED 150.00 15.00 10.00

167 MKT DA 35C CED 21.00 2.80 7.50

168 MKT DA 55C CED 38.20 5.00 7.64

171 CONMACO C 50 ECH 15.00 5.00 3.00

172 CONMACO C 65 ECH 19.50 6.50 3.00

173 CONMACO C 550 ECH 25.00 5.00 5.00

174 CONMACO C 565 ECH 32.50 6.50 5.00

175 CONMACO C 80 ECH 26.00 8.00 3.25

176 CONMACO C 100 ECH 32.50 10.00 3.25

177 CONMACO C 115 ECH 37.38 11.50 3.25

178 CONMACO C 80E5 ECH 40.00 8.00 5.00

179 CONMACO C 100E5 ECH 50.00 10.00 5.00

180 CONMACO C 115E5 ECH 57.50 11.50 5.00

181 CONMACO C 125E5 ECH 62.50 12.50 5.00

182 CONMACO C 140 ECH 42.00 14.00 3.00

183 CONMACO C 160 ECH 48.75 16.25 3.00

184 CONMACO C 200 ECH 60.00 20.00 3.00

185 CONMACO C 300 ECH 90.00 30.00 3.00

186 CONMACO C 5200 ECH 100.00 20.00 5.00

187 CONMACO C 5300 ECH 150.00 30.00 5.00

188 CONMACO C 5450 ECH 225.00 45.00 5.00

189 CONMACO C 5700 ECH 350.00 70.00 5.00

190 CONMACO C 6850 ECH 510.00 85.00 6.00

191 CONMACO C 160 ** ECH 51.78 17.26 3.00

192 CONMACO C 50E5 ECH 25.00 5.00 5.00

193 CONMACO C 65E5 ECH 32.50 6.50 5.00

194 CONMACO C 200E5 ECH 100.00 20.00 5.00

195 CONMACO C 300E5 ECH 150.00 30.00 5.00

196 CONMACO C 1750 ECH 1050.00 175.00 6.00

204 VULCAN VUL 01 ECH 15.00 5.00 3.00

205 VULCAN VUL 02 ECH 7.26 3.00 2.42

206 VULCAN VUL 06 ECH 19.50 6.50 3.00

207 VULCAN VUL 08 ECH 26.00 8.00 3.25

208 VULCAN VUL 010 ECH 32.50 10.00 3.25

209 VULCAN VUL 012 ECH 39.00 12.00 3.25

210 VULCAN VUL 014 ECH 42.00 14.00 3.00

211 VULCAN VUL 016 ECH 48.75 16.25 3.00

212 VULCAN VUL 020 ECH 60.00 20.00 3.00

213 VULCAN VUL 030 ECH 90.00 30.00 3.00

214 VULCAN VUL 040 ECH 120.00 40.00 3.00
523

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

215 VULCAN VUL 060 ECH 180.00 60.00 3.00

220 VULCAN VUL 30C ECH 7.26 3.00 2.42

221 VULCAN VUL 50C ECH 15.10 5.00 3.02

222 VULCAN VUL 65C ECH 19.18 6.50 2.95

223 VULCAN VUL 65CA ECH 19.57 6.50 3.01

224 VULCAN VUL 80C ECH 24.48 8.00 3.06

225 VULCAN VUL 85C ECH 25.99 8.52 3.05

226 VULCAN VUL 100C ECH 32.90 10.00 3.29

227 VULCAN VUL 140C ECH 35.98 14.00 2.57

228 VULCAN VUL 200C ECH 50.20 20.00 2.51

229 VULCAN VUL 400C ECH 113.60 40.00 2.84

230 VULCAN VUL 600C ECH 179.16 60.00 2.99

231 VULCAN VUL 320 ECH 60.00 20.00 3.00

232 VULCAN VUL 330 ECH 90.00 30.00 3.00

233 VULCAN VUL 340 ECH 120.00 40.00 3.00

234 VULCAN VUL 360 ECH 180.00 60.00 3.00

235 VULCAN VUL 505 ECH 25.00 5.00 5.00

236 VULCAN VUL 506 ECH 32.50 6.50 5.00

237 VULCAN VUL 508 ECH 40.00 8.00 5.00

238 VULCAN VUL 510 ECH 50.00 10.00 5.00

239 VULCAN VUL 512 ECH 60.00 12.00 5.00

240 VULCAN VUL 520 ECH 100.00 20.00 5.00

241 VULCAN VUL 530 ECH 150.00 30.00 5.00

242 VULCAN VUL 540 ECH 200.00 40.90 4.89

243 VULCAN VUL 560 ECH 300.00 62.50 4.80

245 VULCAN VUL 3100 ECH 300.00 100.00 3.00

246 VULCAN VUL 5100 ECH 500.00 100.00 5.00

247 VULCAN VUL 5150 ECH 750.00 150.00 5.00

248 VULCAN VUL 6300 ECH 1800.00 300.00 6.00

251 RAYMOND R 1 ECH 15.00 5.00 3.00

252 RAYMOND R 1S ECH 19.50 6.50 3.00

253 RAYMOND R 65C ECH 19.50 6.50 3.00

254 RAYMOND R 65CH ECH 19.50 6.50 3.00

255 RAYMOND R 0 ECH 24.38 7.50 3.25

256 RAYMOND R 80C ECH 24.48 8.00 3.06

257 RAYMOND R 80CH ECH 24.48 8.00 3.06

258 RAYMOND R 2/0 ECH 32.50 10.00 3.25

259 RAYMOND R 3/0 ECH 40.63 12.50 3.25

260 RAYMOND R 150C ECH 48.75 15.00 3.25

261 RAYMOND R 4/0 ECH 48.75 15.00 3.25

262 RAYMOND R 5/0 ECH 56.88 17.50 3.25

263 RAYMOND R 30X ECH 75.00 30.00 2.50

264 RAYMOND R 8/0 ECH 81.25 25.00 3.25
524

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

265 RAYMOND R 40X ECH 100.00 40.00 2.50

266 RAYMOND R 60X ECH 150.00 60.00 2.50

270 MENCK MHU 100C ECH 73.71 11.10 6.64

271 MENCK MH 68 ECH 49.18 7.72 6.37

272 MENCK MH 96 ECH 69.43 11.02 6.30

273 MENCK MH 145 ECH 104.80 16.53 6.34

274 MENCK MHU 195 ECH 143.74 21.36 6.73

275 MENCK MHU 220 ECH 162.17 24.84 6.53

276 MENCK MHU 400 ECH 294.82 51.09 5.77

277 MENCK MHU 600 ECH 442.28 75.52 5.86

278 MENCK MHU 1000 ECH 737.38 126.98 5.81

279 MENCK MHU 1700 ECH 1253.24 207.15 6.05

280 MENCK MHU 2100 ECH 1548.29 257.18 6.02

281 MENCK MHU 3000 ECH 2211.90 370.23 5.97

282 MENCK MRBS 500 ECH 45.07 11.02 4.09

283 MENCK MRBS 750 ECH 67.77 16.53 4.10

285 MENCK MRBS 850 ECH 93.28 18.96 4.92

286 MENCK MRBS1100 ECH 123.43 24.25 5.09

287 MENCK MRBS1502 ECH 135.59 33.07 4.10

288 MENCK MRBS1800 ECH 189.81 38.58 4.92

289 MENCK MRBS2500 ECH 262.11 63.93 4.10

290 MENCK MRBS2502 ECH 225.95 55.11 4.10

291 MENCK MRBS2504 ECH 225.95 55.11 4.10

292 MENCK MRBS3000 ECH 325.36 66.13 4.92

293 MENCK MRBS3900 ECH 513.34 86.86 5.91

294 MENCK MRBS4600 ECH 498.94 101.41 4.92

295 MENCK MRBS5000 ECH 542.33 110.23 4.92

296 MENCK MRBS6000 ECH 759.23 132.27 5.74

297 MENCK MRBS7000 ECH 631.40 154.00 4.10

298 MENCK MRBS8000 ECH 867.74 176.37 4.92

299 MENCK MRBS8800 ECH 954.53 194.01 4.92

300 MENCK MBS12500 ECH 1581.83 275.58 5.74

301 MKT No. 5 ECH 1.00 0.20 5.00

302 MKT No. 6 ECH 2.50 0.40 6.25

303 MKT No. 7 ECH 4.15 0.80 5.19

304 MKT 9B3 ECH 8.75 1.60 5.47

305 MKT 10B3 ECH 13.11 3.00 4.37

306 MKT C5-Air ECH 14.20 5.00 2.84

307 MKT C5-Steam ECH 16.20 5.00 3.24

308 MKT S-5 ECH 16.25 5.00 3.25

309 MKT 11B3 ECH 19.15 5.00 3.83

310 MKT C826 Stm ECH 24.40 8.00 3.05

311 MKT C826 Air ECH 21.20 8.00 2.65
525

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

312 MKT S-8 ECH 26.00 8.00 3.25

313 MKT MS-350 ECH 30.80 7.72 3.99

314 MKT S 10 ECH 32.50 10.00 3.25

315 MKT S 14 ECH 37.52 14.00 2.68

316 MKT MS 500 ECH 44.00 11.00 4.00

317 MKT S 20 ECH 60.00 20.00 3.00

318 IHC S-30 ECH 21.70 3.53 6.15

319 IHC S-40 ECH 28.93 4.85 5.97

320 IHC S-35 ECH 25.53 6.63 3.85

321 IHC S-70 ECH 51.25 7.73 6.63

322 IHC S-90 ECH 65.90 9.94 6.63

323 IHC S-120 ECH 89.37 13.48 6.63

324 IHC S-150 ECH 110.06 16.60 6.63

325 IHC S-200 ECH 145.64 22.00 6.62

326 IHC S-280 ECH 205.31 30.06 6.83

327 IHC S-400 ECH 292.60 44.20 6.62

328 IHC S-500 ECH 366.09 55.30 6.62

329 IHC S-600 ECH 443.54 67.00 6.62

330 IHC S-900 ECH 658.36 99.45 6.62

331 IHC S-1200 ECH 891.05 134.60 6.62

332 IHC S-1800-L ECH 1170.39 166.00 7.05

333 IHC S-2300 ECH 1681.48 254.00 6.62

334 IHC S-2000 ECH 1473.97 222.65 6.62

335 IHC SC-30 ECH 21.81 3.76 5.80

336 IHC SC-40 ECH 29.86 5.51 5.42

337 IHC SC-50 ECH 36.82 7.29 5.05

338 IHC SC-60 ECH 44.95 13.30 3.38

339 IHC SC-75 ECH 54.80 12.15 4.51

340 IHC SC-110 ECH 81.89 17.46 4.69

341 IHC SC-150 ECH 109.35 24.30 4.50

342 IHC SC-200 ECH 152.51 30.20 5.05

343 IHC SC-250 ECH 179.80 37.26 4.83

344 IHC S-750 ECH 550.79 83.11 6.63

345 IHC S-800 ECH 589.97 88.15 6.69

346 IHC S-1400 ECH 1033.84 147.94 6.99

347 IHC S-1800 ECH 1340.21 195.64 6.85

348 IHC S-2500 ECH 1843.16 275.80 6.68

349 HERA 1900 OED 44.41 4.19 10.60

350 HERA 1250 OED 24.85 2.76 9.02

351 HERA 1500 OED 29.81 3.31 9.02

352 HERA 2500 OED 49.70 5.51 9.02

353 HERA 2800 OED 55.70 6.18 9.02

354 HERA 3500 OED 69.59 7.72 9.02
526

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

355 HERA 5000 OED 99.45 11.03 9.02

356 HERA 5700 OED 113.38 12.57 9.02

357 HERA 6200 OED 123.30 13.67 9.02

358 HERA 7500 OED 149.19 16.54 9.02

359 HERA 8800 OED 174.99 19.40 9.02

360 ICE I-12obs OED 30.21 2.82 10.71

361 ICE I-19obs OED 43.24 4.02 10.77

362 ICE I-30obs OED 71.45 6.62 10.80

363 ICE I-36obs OED 90.68 7.94 11.42

364 ICE I-46obs OED 107.74 10.15 10.62

365 ICE I-62obs OED 164.98 14.60 11.30

366 ICE I-80obs OED 212.40 17.70 12.00

367 ICE I-8v2obs OED 17.60 1.76 10.00

368 ICE I-100obs OED 264.45 23.61 11.20

369 BSP SL20 ECH 14.11 3.31 4.27

370 BSP SL30 ECH 21.69 5.51 3.94

371 FAIRCHLD F-45 ECH 45.00 15.00 3.00

372 FAIRCHLD F-32 ECH 32.55 10.85 3.00

374 BSP CX40 ECH 28.21 6.61 4.27

375 BSP CX50 ECH 37.61 8.82 4.27

376 BSP CX60 ECH 47.01 11.02 4.27

377 BSP CX75 ECH 52.08 13.23 3.94

378 BSP CX85 ECH 60.75 15.43 3.94

379 BSP CX110 ECH 78.11 19.84 3.94 

381 BSP HH3 ECH 26.02 6.61 3.94

382 BSP HH5 ECH 43.38 11.02 3.94

383 BSP HH7 ECH 60.78 15.43 3.94

384 BSP HH8 ECH 69.50 17.64 3.94

385 BSP HH9 ECH 78.17 19.84 3.94

386 BSP HH11-1.2 ECH 95.55 24.25 3.94

387 BSP HH14-1.2 ECH 121.59 30.86 3.94

388 BSP HH16-1.2 ECH 138.87 35.27 3.94

391 BSP HA30 ECH 260.37 66.14 3.94

392 BSP HA40 ECH 347.16 88.18 3.94

393 BSP HH11-1.5 ECH 119.31 24.25 4.92

394 BSP HH14-1.5 ECH 151.83 30.86 4.92

395 BSP HH16-1.5 ECH 173.54 35.27 4.92

396 BSP CG180 ECH 131.92 26.45 4.99

397 BSP CG210 ECH 153.91 30.86 4.99

398 BSP CG240 ECH 175.90 35.27 4.99

399 BSP CG270 ECH 197.88 39.68 4.99

400 BSP CG300 ECH 219.87 44.09 4.99

401 BERMINGH B23 CED 22.99 2.80 8.21
527

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

402 BERMINGH B200 OED 18.00 2.00 9.00

403 BERMINGH B225 OED 29.25 3.00 9.75

404 BERMINGH B300 OED 40.31 3.75 10.75

405 BERMINGH B400 OED 53.75 5.00 10.75

406 BERMINGH B-21 OED 53.25 4.63 11.50

410 BERMINGH B300 M OED 40.31 3.75 10.75

411 BERMINGH B400 M OED 53.75 5.00 10.75

412 BERMINGH B400 4.8 OED 43.20 4.80 9.00

413 BERMINGH B400 5.0 OED 45.00 5.00 9.00

414 BERMINGH B23 5 CED 22.99 2.80 8.21

415 BERMINGH B250 5 OED 26.25 2.50 10.50

416 BERMINGH B3505 OED 47.20 4.00 11.80

417 BERMINGH B4005 OED 59.00 5.00 11.80

418 BERMINGH B4505 OED 77.88 6.60 11.80

419 BERMINGH B5005 OED 92.04 7.80 11.80

420 BERMINGH B5505 OED 108.56 9.20 11.80

421 BERMINGH B550 C OED 88.00 11.00 8.00

422 BERMINGH B2005 OED 18.00 2.00 9.00

424 BERMINGH B2505 OED 35.40 3.00 11.80

425 BERMINGH B3005 OED 35.40 3.00 11.80

431 BERMINGH B6005 OED 160.95 13.64 11.80

432 BERMINGH B6505 C OED 253.00 22.00 11.50

433 BERMINGH B6505 OED 202.86 17.64 11.50

434 BERMINGH B-9 OED 21.00 2.00 10.50

435 BERMINGH B-32 OED 81.08 7.05 11.50

436 BERMINGH B-64 OED 166.50 14.11 11.80

437 BERMINGH B-6505HD OED 220.50 22.05 10.00

441 MENCK MHF5-5 ECH 38.69 11.02 3.51

442 MENCK MHF5-6 ECH 46.43 13.23 3.51

443 MENCK MHF5-7 ECH 54.17 15.43 3.51

444 MENCK MHF5-8 ECH 61.91 17.64 3.51

445 MENCK MHF5-9 ECH 69.65 19.84 3.51

446 MENCK MHF5-10 ECH 77.39 22.05 3.51

447 MENCK MHF5-11 ECH 85.13 24.25 3.51

448 MENCK MHF5-12 ECH 92.87 26.45 3.51

449 MENCK MHF3-3 ECH 24.76 7.05 3.51

450 MENCK MHF3-4 ECH 30.96 8.82 3.51

451 MENCK MHF3-5 ECH 38.69 11.02 3.51

452 MENCK MHF3-6 ECH 46.43 13.23 3.51

453 MENCK MHF3-7 ECH 54.17 15.43 3.51

454 MENCK MHF10-15 ECH 124.73 33.06 3.77

455 MENCK MHF10-20 ECH 166.28 44.07 3.77

456 MENCK MHF 5-14 ECH 108.34 30.86 3.51
528

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

457 MENCK MHU135T* ECH 110.59 17.99 6.15

458 MENCK MHU500T* ECH 368.74 65.96 5.59

459 MENCK MHU 300S ECH 221.20 35.73 6.19

460 MENCK MHU 270T ECH 221.20 35.73 6.19

461 MENCK MHU 200T ECH 162.24 26.75 6.07

462 MENCK MHU 400T ECH 324.37 52.45 6.18

463 MENCK MHU 500T ECH 405.53 65.96 6.15

464 MENCK MHU 700T ECH 567.72 92.88 6.11

465 MENCK MHU 840S ECH 619.22 92.88 6.67

466 MENCK MHU 600B ECH 457.03 65.96 6.93

467 MENCK MHU 600T ECH 486.63 80.39 6.05

468 MENCK MHU 800S ECH 604.57 99.93 6.05

469 MENCK MHU1200S ECH 884.84 145.71 6.07

470 MENCK MHU1500S ECH 1106.07 178.94 6.18

471 MENCK MHU1700T ECH 1400.86 227.36 6.16

472 MENCK MHU1900S ECH 1400.86 227.36 6.16

473 MENCK MHU 150S ECH 110.59 17.99 6.15

474 MENCK MHU2700S ECH 1990.19 318.77 6.24

475 MENCK MHU 135T ECH 110.59 17.99 6.15

476 MENCK MHU 750T ECH 604.57 99.93 6.05

477 MENCK MHU1100T ECH 899.66 145.71 6.18

478 MENCK MHU150S* ECH 110.59 17.99 6.15

479 MENCK MHU600B* ECH 457.03 65.96 6.93

481 JUNTTAN HHK3A ECH 26.05 6.62 3.94

482 JUNTTAN HHK4A ECH 34.73 8.82 3.94

483 JUNTTAN HHK5A ECH 43.41 11.03 3.94

484 JUNTTAN HHK6A ECH 52.10 13.23 3.94

485 JUNTTAN HHK7A ECH 60.75 15.43 3.94

486 JUNTTAN HHK10A ECH 86.83 22.05 3.94

487 JUNTTAN HHK12A ECH 104.19 26.47 3.94

488 JUNTTAN HHK14A ECH 121.56 30.88 3.94

491 JUNTTAN HHK9A ECH 78.14 19.85 3.94

494 JUNTTAN HHK16A ECH 138.92 35.29 3.94

495 JUNTTAN HHK18A ECH 156.29 39.70 3.94

496 JUNTTAN HHK20A ECH 173.65 44.11 3.94

497 JUNTTAN HHK4SL ECH 43.40 8.82 4.92

498 JUNTTAN HHK3AL ECH 17.37 6.62 2.63

499 JUNTTAN HHK4AL ECH 23.15 8.82 2.63

500 JUNTTAN HHK5AL ECH 28.94 11.03 2.63

501 HPSI 110 ECH 44.00 11.00 4.00

502 HPSI 150 ECH 60.00 15.00 4.00

503 HPSI 154 ECH 61.60 15.40 4.00

504 HPSI 200 ECH 80.00 20.00 4.00
529

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

505 HPSI 225 ECH 90.00 22.50 4.00

506 HPSI 650 ECH 32.50 6.50 5.00

507 HPSI 1000 ECH 50.00 10.00 5.00

508 HPSI 1605 ECH 83.00 16.60 5.00

509 HPSI 2005 ECH 95.10 19.02 5.00

510 HPSI 3005 ECH 154.33 30.87 5.00

511 HPSI 3505 ECH 176.33 35.27 5.00

512 HPSI 2000 ECH 80.00 20.00 4.00

514 UDDCOMB H2H ECH 16.62 4.40 3.77

515 UDDCOMB H3H ECH 24.88 6.60 3.77

516 UDDCOMB H4H ECH 33.18 8.80 3.77

517 UDDCOMB H5H ECH 41.47 11.00 3.77

518 UDDCOMB H6H ECH 49.76 13.20 3.77

519 UDDCOMB H8H ECH 82.19 17.60 4.67

520 UDDCOMB H10H ECH 86.88 22.05 3.94

521 DAWSON HPH1200 ECH 8.72 2.30 3.79

522 DAWSON HPH1800 ECH 13.72 3.30 4.16

523 DAWSON HPH2400 ECH 17.32 4.19 4.13

524 DAWSON HPH6500 ECH 46.98 10.25 4.58

525 DAWSON HPH4500 ECH 32.56 7.72 4.22

526 DAWSON HPH9000 ECH 66.30 10.47 6.33

530 BRUCE SGH-0312 ECH 26.00 6.60 3.94

531 BRUCE SGH-0512 ECH 43.34 11.00 3.94

532 BRUCE SGH-0712 ECH 60.68 15.40 3.94

533 BRUCE SGH-1012 ECH 86.77 22.05 3.94

534 BRUCE SGH-0412 ECH 34.67 8.80 3.94

535 BANUT S3000 ECH 26.04 6.62 3.94

536 BANUT S4000 ECH 34.72 8.82 3.94

537 BANUT S5000 ECH 43.41 11.03 3.94

538 BANUT S6000 ECH 52.09 13.23 3.94

539 BANUT S8000 ECH 69.45 17.64 3.94

540 BANUT S10000 ECH 86.81 22.05 3.94

541 BANUT 3 Tonnes ECH 17.35 6.61 2.62

542 BANUT 4 Tonnes ECH 23.14 8.82 2.62

543 BANUT 5 Tonnes ECH 28.92 11.02 2.62

544 BANUT 6 Tonnes ECH 34.72 13.23 2.62

545 BANUT 7 Tonnes ECH 40.49 15.43 2.62

550 ICE 70 ECH 21.00 7.00 3.00

551 ICE 75 ECH 30.00 7.50 4.00

552 ICE 110-SH ECH 37.72 11.50 3.28

553 ICE 115-SH ECH 37.95 11.50 3.30

554 ICE 115 ECH 46.00 11.50 4.00

555 ICE 160-SH ECH 64.00 16.00 4.00
530

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

556 ICE 160 ECH 64.00 16.00 4.00

557 ICE 220 ECH 88.00 22.00 4.00

558 ICE 275 ECH 110.00 27.50 4.00

559 ICE DKH-3U ECH 26.00 6.60 3.94

560 HMC 28A ECH 28.00 7.00 4.00

561 HMC 28B ECH 21.00 7.00 3.00

562 HMC 62 ECH 46.00 11.50 4.00

563 HMC 86 ECH 64.00 16.00 4.00

564 HMC 119 ECH 88.00 22.00 4.00

565 HMC 149 ECH 110.00 27.50 4.00

566 HMC 187 ECH 138.00 34.50 4.00

567 HMC 19D ECH 14.00 3.50 4.00

568 HMC 38D ECH 28.00 7.00 4.00

569 APE D 8-42 OED 19.80 1.76 11.25

570 APE D 1-42 OED 1.32 0.21 6.33

571 APE D 19-42 OED 47.13 4.19 11.25

572 APE D 30-42 OED 74.42 6.62 11.25

573 APE D 36-42 OED 89.30 7.94 11.25

574 APE D 46-42 OED 114.11 10.14 11.25

575 APE D 62-42 OED 153.80 13.67 11.25

576 APE D 80-42 OED 198.45 17.64 11.25

577 APE D 100-42 OED 248.06 22.05 11.25

579 APE D 16-42 OED 39.69 3.53 11.25

580 APE D 16-52 OED 39.69 3.53 11.25

581 APE D 25-42 OED 62.01 5.51 11.25

582 APE D 125-42 OED 310.08 27.56 11.25

583 APE D 50-42 OED 124.03 11.03 11.25

584 APE D 12-42 OED 29.77 2.65 11.25

585 APE D 36-26 OED 89.30 7.94 11.25

586 APE D 128-42 OED 317.25 28.20 11.25

587 APE D 138-42 OED 342.00 30.40 11.25

588 APE D 160-42 OED 396.90 35.28 11.25

589 APE D 180-42 OED 446.51 39.69 11.25

590 APE D 225-42 OED 558.00 49.60 11.25

591 APE 5.4mT ECH 26.00 12.00 2.17

592 APE 7.2mT ECH 51.30 16.20 3.17

593 APE D 220-42 OED 540.81 48.46 11.16

594 APE 15-60 ECH 150.00 30.00 5.00

595 APE 10-60 ECH 100.00 20.00 5.00

596 APE 400U ECH 400.00 80.00 5.00

598 APE 750U ECH 750.00 120.00 6.25

599 APE D 100-13 OED 300.04 23.70 12.66

600 BSP DX20 ECH 14.11 3.31 4.27
531

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

601 BSP DX25 ECH 18.09 4.41 4.10

602 BSP DX30 ECH 21.71 5.51 3.94

603 BSP LX2.5-SA ECH 14.47 5.51 2.63

604 BSP LX4-SA ECH 23.15 8.82 2.63

605 BSP LX5-SA ECH 28.94 11.03 2.63

606 BSP CGL370 ECH 271.22 55.11 4.92

607 BSP CGL440 ECH 325.47 66.14 4.92

608 BSP CGL520 ECH 379.71 77.16 4.92

609 BSP CGL590 ECH 433.96 88.18 4.92

610 BSP LX7-SA ECH 40.52 15.44 2.63

625 BRUCE SGH-1212 ECH 104.13 26.46 3.94

626 BRUCE SGH-1312 ECH 112.81 28.66 3.94

627 BRUCE SGH-1315 ECH 141.01 28.66 4.92

628 BRUCE SGH-1412 ECH 121.48 30.87 3.94

629 BRUCE SGH-1415 ECH 151.85 30.87 4.92

630 BRUCE SGH-1612 ECH 138.84 35.27 3.94

631 BRUCE SGH-1615 ECH 173.55 35.27 4.92

632 BRUCE SGH-1618 ECH 208.26 35.27 5.90

633 BRUCE SGH-1619 ECH 219.83 35.27 6.23

634 BRUCE SGH-1812 ECH 156.19 39.68 3.94

635 BRUCE SGH-1815 ECH 195.24 39.68 4.92

636 BRUCE SGH-2012 ECH 173.55 44.09 3.94

637 BRUCE SGH-2015 ECH 216.94 44.09 4.92

638 BRUCE SGH-2312 ECH 199.58 50.71 3.94

639 BRUCE SGH-2315 ECH 249.48 50.71 4.92

640 BRUCE SGH-3012 ECH 260.32 66.14 3.94

641 BRUCE SGH-3013 ECH 282.02 66.14 4.26

642 BRUCE SGH-3015 ECH 325.40 66.14 4.92

643 BRUCE SGH-4012 ECH 347.10 88.19 3.94

644 BRUCE SGH-4212 ECH 364.45 92.59 3.94

645 BRUCE SGH-5012 ECH 433.87 110.23 3.94

650 Twinwood V20B ECH 35.58 9.04 3.94

651 Twinwood V100D ECH 87.66 22.27 3.94

652 Twinwood V160B ECH 140.58 35.71 3.94

653 Twinwood V400A ECH 263.84 67.02 3.94

656 Pilemast 24-750 ECH 1.50 0.75 2.00

657 Pilemast 24-900 ECH 1.80 0.90 2.00

658 Pilemast 24-2000 ECH 4.00 2.00 2.00

659 Pilemast 24-2500 ECH 5.00 2.50 2.00

660 Pilemast 36-3000 ECH 9.00 3.00 3.00

661 Pilemast 36-5000 ECH 15.00 5.00 3.00

669 MVE M-12 OED 30.21 2.82 10.71

670 MVE M-19 OED 49.38 4.02 12.30
532

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

671 MVE M-30 OED 83.35 6.62 12.60

801 DKH PH-5 ECH 43.40 11.02 3.94

802 DKH PH-7 ECH 60.75 15.43 3.94

803 DKH PH-7S ECH 60.75 15.43 3.94

804 DKH PH-10 ECH 86.79 22.05 3.94

805 DKH PH-13 ECH 112.83 28.66 3.94

806 DKH PH-20 ECH 216.98 44.09 4.92

807 DKH PH-30 ECH 325.47 66.14 4.92 

808 DKH PH-40 ECH 433.96 88.18 4.92

809 DKH DKH-713 ECH 112.92 28.66 3.94

850 PILECO D8-22 OED 18.66 1.76 10.60

851 PILECO D12-42 OED 29.89 2.82 10.60

852 PILECO D19-42 OED 42.51 4.01 10.60

853 PILECO D25-32 OED 58.41 5.51 10.60

854 PILECO D30-32 OED 70.07 6.61 10.60

855 PILECO D36-32 OED 84.16 7.94 10.60

856 PILECO D46-32 OED 107.48 10.14 10.60

857 PILECO D62-22 OED 161.31 13.67 11.80

858 PILECO D80-23 OED 197.57 17.64 11.20

859 PILECO D100-13 OED 246.85 22.04 11.20

860 PILECO D125-32 OED 308.67 27.56 11.20

861 PILECO D225-22 OED 555.34 49.58 11.20

862 PILECO D250-22 OED 617.06 55.09 11.20

863 PILECO D138-32 OED 340.61 30.41 11.20

864 PILECO D180-32 OED 444.27 39.67 11.20

865 PILECO D280-22 OED 688.55 61.73 11.16

866 PILECO D160-32 OED 395.08 35.28 11.20

867 PILECO D400-12 OED 810.10 88.15 9.19

868 PILECO D600-12 OED 1215.10 132.22 9.19

869 PILECO D800-22 OED 1620.20 176.30 9.19

921 BRUCE SGH-0212 ECH 17.34 4.40 3.94

922 BRUCE SGH-0715 ECH 75.77 15.40 4.92

923 BRUCE SGH-1015 ECH 108.47 22.05 4.92

924 BRUCE SGH-1215 ECH 130.16 26.46 4.92

925 BRUCE SGH-2512 ECH 216.94 55.12 3.94

926 BRUCE SGH-2515 ECH 271.17 55.12 4.92

927 BRUCE SGH-3512 ECH 303.71 77.16 3.94

928 BRUCE SGH-3515 ECH 379.64 77.16 4.92

929 BRUCE SGH-4015 ECH 433.87 88.19 4.92

930 BRUCE SGH-4215 ECH 455.56 92.59 4.92

931 BRUCE SGH-4512 ECH 390.48 99.21 3.94

932 BRUCE SGH-4515 ECH 488.10 99.21 4.92

933 BRUCE SGH-4712 ECH 407.84 103.62 3.94
533

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

934 BRUCE SGH-4715 ECH 509.80 103.62 4.92

935 BRUCE SGH-4719 ECH 645.74 103.62 6.23

936 BRUCE SGH-5015 ECH 542.34 110.23 4.92

937 BRUCE SGH-5715 ECH 618.26 125.66 4.92

938 BRUCE SGH-6015 ECH 650.80 132.28 4.92

939 BRUCE SGH-7015 ECH 759.27 154.32 4.92

940 BRUCE SGH-8015 ECH 867.74 176.37 4.92

949 JUNTTAN HHK10S ECH 108.49 22.05 4.92

950 JUNTTAN HHK28S ECH 303.78 61.73 4.92

951 JUNTTAN HHK5S ECH 54.27 11.03 4.92

952 JUNTTAN HHK7S ECH 75.97 15.44 4.92

953 JUNTTAN HHK9S ECH 97.68 19.85 4.92

954 JUNTTAN HHK12S ECH 130.24 26.47 4.92

955 JUNTTAN HHK14S ECH 151.95 30.88 4.92

956 JUNTTAN HHK16S ECH 173.65 35.29 4.92

957 JUNTTAN HHK18S ECH 195.36 39.70 4.92

958 JUNTTAN HHK20S ECH 217.07 44.11 4.92

959 JUNTTAN HHK25S ECH 271.22 55.11 4.92

960 JUNTTAN HHK36S ECH 390.56 79.36 4.92

961 JUNTTAN HHU5A ECH 54.27 11.03 4.92

962 JUNTTAN HHU7A ECH 75.94 15.43 4.92

963 JUNTTAN HHU9A ECH 97.64 19.84 4.92

964 JUNTTAN HHU12A ECH 130.19 26.45 4.92

965 JUNTTAN HHU14A ECH 151.88 30.86 4.92

966 JUNTTAN HHU16A ECH 173.58 35.27 4.92

968 JUNTTAN SHK100-3 ECH 26.91 6.61 4.07

969 JUNTTAN SHK100-3 ECH 35.89 8.82 4.07

970 JUNTTAN SHK100-3 ECH 44.84 11.02 4.07

971 JUNTTAN SHK100-3 ECH 53.82 13.23 4.07

972 JUNTTAN SHK110-5 ECH 44.98 11.02 4.08

973 JUNTTAN SHK110-5 ECH 53.82 13.23 4.07

974 JUNTTAN SHK110-5 ECH 65.62 15.43 4.25

975 JUNTTAN SHK110-5 ECH 77.42 17.64 4.39

976 JUNTTAN SHK110-5 ECH 87.74 19.84 4.42

977 JUNTTAN SHK100-5 ECH 44.84 11.02 4.07

978 JUNTTAN SHK100-5 ECH 53.82 13.23 4.07

979 JUNTTAN SHK110-7 ECH 65.63 15.43 4.25

980 JUNTTAN SHK110-7 ECH 77.43 17.64 4.39

981 JUNTTAN SHK110-7 ECH 87.74 19.84 4.42

998 HYPOTHET EX 4 OED 23.38 2.75 8.50

999 SELF Drop/10t ECH 300.00 20.00 15.00

1001 DFI-Corp HHA250-4 ECH 25.18 5.51 4.57

1002 DFI-Corp HHA300-4 ECH 28.75 6.61 4.35
534

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

1003 DFI-Corp HHA325-4 ECH 30.36 7.16 4.24

1004 DFI-Corp HHA350-4 ECH 31.80 7.70 4.13

1005 DFI-Corp HHA400-6 ECH 51.92 8.80 5.90

1006 DFI-Corp HHA450-6 ECH 57.04 9.92 5.75

1007 DFI-Corp HHB500-6 ECH 66.66 11.00 6.06

1008 DFI-Corp HHB600-6 ECH 77.88 13.20 5.90

1020 J&M 70B HIH ECH 21.00 7.00 3.00

1021 J&M 82 HIH ECH 32.80 8.20 4.00

1022 J&M 115 HIH ECH 46.00 11.50 4.00

1023 J&M 160 HIH ECH 64.00 16.00 4.00

1024 J&M 220 HIH ECH 88.00 22.00 4.00

1025 J&M 275 HIH ECH 110.00 27.50 4.00

1026 J&M 345 HIH ECH 138.00 34.50 4.00

1134 Pilemer DKH-3U ECH 26.00 6.60 3.94

1135 Pilemer DKH 10L ECH 86.79 22.05 3.94

1201 Liebherr H 50/3 ECH 28.97 6.60 4.39

1202 Liebherr H 50/4 ECH 35.02 8.80 3.98

1203 Liebherr H 85/5 ECH 43.34 11.00 3.94

1204 Liebherr H 85/7 ECH 60.16 15.43 3.90

1205 Liebherr H 110/7 ECH 60.16 15.43 3.90

1206 Liebherr H 110/9 ECH 78.01 19.85 3.93

1251 ICE I-30 V2 OED 71.71 6.62 10.84

1261 APE D 19-52 OED 47.13 4.19 11.25

1262 APE D 16-32 OED 39.69 3.53 11.25

1263 APE D 19-32 OED 47.13 4.19 11.25

1264 APE D 25-32 OED 62.01 5.51 11.25

1265 APE D 30-32 OED 74.42 6.62 11.25

1266 APE D 36-32 OED 89.30 7.94 11.25

1267 APE D 46-32 OED 114.11 10.14 11.25

1268 APE D 62-22 OED 153.80 13.67 11.25

1269 APE D 80-23 OED 198.45 17.64 11.25

1270 APE D 100-32 OED 248.06 22.05 11.25

1271 APE D 120-32 OED 349.69 27.60 12.67

1272 APE D 70-42 OED 173.64 15.44 11.25

1273 APE D 25-52 OED 62.01 5.51 11.25

1274 APE D 30-52 OED 74.42 6.62 11.25

1275 APE D 36-52 OED 89.30 7.94 11.25

1276 APE D 46-52 OED 114.11 10.14 11.25

1277 APE D 50-52 OED 124.03 11.03 11.25

1278 APE D 62-52 OED 153.80 13.67 11.25

1279 APE D 70-52 OED 173.64 15.44 11.25

1280 APE 7.5a ECH 24.00 12.00 2.00

1281 APE 7.5b ECH 20.40 10.20 2.00
535

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

1282 APE 7.5c ECH 15.20 7.60 2.00

1283 APE 9.5a ECH 50.66 16.00 3.17

1284 APE 9.5b ECH 44.32 14.00 3.17

1285 APE 7-3 ECH 42.00 14.00 3.00

1286 APE 8-3 ECH 48.00 16.00 3.00

1287 APE 8 ECH 16.00 8.00 2.00

1288 APE 8a ECH 24.00 12.00 2.00

1289 APE 10-4 ECH 80.00 20.00 4.00

1321 MENCK MHU 240U ECH 221.20 35.73 6.19

1322 MENCK MHU 440S ECH 324.37 52.45 6.18

1323 MENCK MHU 360U ECH 324.37 52.45 6.18

1324 MENCK MHU 550S ECH 404.06 65.96 6.13

1325 MENCK MHU 450U ECH 404.06 65.96 6.13

1326 MENCK MHU 660S ECH 485.17 80.39 6.04

1327 MENCK MHU 540U ECH 485.17 80.39 6.04

1328 MENCK MHU 720T ECH 588.19 99.93 5.89

1329 MENCK MHU 650U ECH 588.19 99.93 5.89

1330 MENCK MHU1000S ECH 736.48 126.98 5.80

1331 MENCK MHU 900T ECH 736.48 126.98 5.80

1332 MENCK MHU 810U ECH 736.48 126.98 5.80

1333 MENCK MHU1700S ECH 1272.95 207.15 6.15

1334 MENCK MHU2100S ECH 1573.92 257.18 6.12

1335 MENCK MHU3000S ECH 2216.56 370.23 5.99

1336 MENCK MHU1400B ECH 1032.08 145.71 7.08

1337 MENCK MHU3500S ECH 2582.43 385.85 6.69

1371 IHC S-3000 ECH 2211.93 332.44 6.65

1372 IHC S-4000 ECH 2948.91 444.30 6.64

1401 FAMBO HR250 ECH 1.81 0.55 3.28

1402 FAMBO HR500akk ECH 3.62 1.10 3.28

1403 FAMBO HR500 ECH 4.34 1.10 3.94

1404 FAMBO HR1000 ECH 8.68 2.20 3.94

1405 FAMBO HR1500 ECH 13.02 3.31 3.94

1406 FAMBO HR2000 ECH 17.36 4.41 3.94

1407 FAMBO HR2750 ECH 23.87 6.06 3.94

1408 FAMBO HR3000 ECH 26.04 6.61 3.94

1409 FAMBO HR4000 ECH 34.72 8.82 3.94

1410 FAMBO HR5000 ECH 43.40 11.02 3.94

1411 FAMBO HR7000 ECH 60.75 15.43 3.94

1412 FAMBO HR8000 ECH 69.45 17.64 3.94

1413 FAMBO HR10000 ECH 86.79 22.05 3.94

1501 ICE I-12v2 OED 29.63 2.82 10.50

1502 ICE I-8v2 OED 18.69 1.76 10.60

1503 ICE I-19v2 OED 46.14 4.01 11.50
536

Type Hammer 

Make

Hammer 

Model

Hammer 

Type

Rated Energy 

kip-feet

Ram Weight 

kips

Stroke 

feet

1504 ICE I-30v2 OED 76.05 6.61 11.50

1505 ICE I-36v2 OED 93.73 7.94 11.81

1506 ICE I-46v2 OED 119.77 10.14 11.81

1507 ICE I-62v2 OED 172.37 14.59 11.81

1508 ICE I-80v2 OED 208.30 17.64 11.81

1509 ICE I-100v2 OED 260.37 22.05 11.81

1510 ICE I-125v2 OED 310.10 27.56 11.25

1511 ICE I-160v2 OED 393.45 35.27 11.16

1512 ICE IP-2 ECH 17.36 4.41 3.94

1513 ICE IP-3 ECH 26.04 6.61 3.94

1514 ICE IP-5 ECH 43.40 11.02 3.94

1515 ICE IP-7 ECH 60.75 15.43 3.94

1516 ICE IP-10 ECH 86.78 22.04 3.94

1517 ICE IP-13 ECH 112.83 28.66 3.94

1520 ICE I-138v2 OED 328.62 30.40 10.81

1531 SPI D 19-42 OED 42.61 4.02 10.60

1532 SPI D 30-32 OED 72.08 6.80 10.60

1601 DELMAG D 2 OED 1.78 0.49 3.61

1602 DELMAG D 4 OED 3.60 0.84 4.30

1603 DELMAG D 8-12 OED 20.10 1.76 11.42

1604 DELMAG D 12-52 OED 33.98 2.82 12.05

1605 DELMAG D 16-52 OED 40.20 3.52 11.42

1606 DELMAG D 25-52 OED 66.34 5.51 12.04

1607 DELMAG D 30-52 OED 75.44 6.60 11.43

1611 DELMAG D138-32 OED 339.51 30.44 11.16

1612 DELMAG D180-32 OED 442.64 39.68 11.16

1613 DELMAG D300-32 OED 737.73 66.14 11.16

1614 DELMAG D400-32 OED 983.64 88.18 11.16

1620 HMC TD19 OED 46.09 4.01 11.49

1621 HMC TD30 OED 69.87 6.61 10.57
537

Type Hammer 

Make 

Hammer 

Model

Hammer 

Type

Rated Power

kW 

Ecc. Mass

kips 

Frequency

Hz 

620 MAIT 34 VIB 227.00 1.23 33.30

621 MAIT 42 VIB 309.00 1.52 33.30

622 MAIT 54 VIB 450.00 0.98 33.30

623 MAIT 68 VIB 531.00 1.23 33.30

624 MAIT 120 VIB 674.00 1.74 30.00

698 ICE 50B VIB 432.00 10.42 26.70

699 ICE 3117 VIB 235.00 1.12 28.30

700 ICE 23-28 VIB 21.00 0.10 26.70

701 ICE 216 VIB 130.00 0.46 26.70

702 ICE 216E VIB 130.00 0.46 26.70

703 ICE 23-Nov VIB 164.00 0.46 31.70

704 ICE 223 VIB 242.00 0.46 38.30

705 ICE 416L VIB 242.00 0.92 26.70

706 ICE 812 VIB 375.00 1.82 26.70

707 ICE 815 VIB 375.00 1.84 26.70

708 ICE 44-30 VIB 242.00 1.30 20.00

709 ICE 44-50 VIB 377.00 1.30 26.70

710 ICE 44-65 VIB 485.00 1.30 27.50

711 ICE 66-65 VIB 485.00 1.95 21.70

712 ICE 66-80 VIB 597.00 1.95 26.70

713 ICE 1412B VIB 597.00 2.04 21.00

714 ICE 1412C VIB 470.00 2.02 23.00

715 ICE V125 VIB 984.00 1.04 25.80

716 ICE 14RF VIB 242.00 1.01 38.30

717 ICE 14-23 VIB 164.00 1.17 35.00

718 ICE 22-23V VIB 164.00 0.92 26.90

719 ICE 22-30 VIB 250.00 0.92 26.90

720 HMC 3+28 VIB 21.00 0.11 26.80

721 HMC 3+75 VIB 56.00 0.11 36.10

722 HMC 13+200 VIB 149.00 0.35 26.70

723 HMC 13S+200 VIB 149.00 0.35 26.70

724 HMC 13H+200 VIB 164.00 0.35 29.80

725 HMC 25+220 VIB 164.00 0.61 20.90

726 HMC 26+335 VIB 242.00 0.71 25.60

727 HMC 26S+335 VIB 242.00 0.71 25.60

728 HMC 51+335 VIB 242.00 1.21 19.50

729 HMC 51+535 VIB 377.00 1.21 26.40

730 HMC 51S+535 VIB 377.00 1.21 26.40

731 HMC 51+740 VIB 485.00 1.21 27.50

732 HMC 76+740 VIB 485.00 1.82 21.70

733 HMC 76+800 VIB 597.00 1.82 26.10

734 HMC 115+800 VIB 597.00 1.35 20.40

735 HMC 230+1600 VIB 1193.00 2.69 20.40
538

Type Hammer 

Make 

Hammer 

Model

Hammer 

Type

Rated Power

kW 

Ecc. Mass

kips 

Frequency

Hz 

750 MKT V-2B VIB 52.00 0.15 30.00

751 MKT V-5C VIB 138.00 0.43 28.33

752 MKT V-20B VIB 242.00 0.20 28.33

753 MKT V-30 VIB 448.00 1.47 28.33

754 MKT V-35 VIB 485.00 1.60 28.33

755 MKT V-140 VIB 1341.00 1.17 23.33

770 APE 3 VIB 10.58 0.00 38.30

771 APE 6 VIB 10.58 0.01 38.30

772 APE 15 VIB 59.67 0.11 30.00

773 APE 20 VIB 59.67 0.15 38.30

774 APE 20E VIB 59.67 0.15 38.30

775 APE 50 VIB 194.00 0.23 30.00

776 APE 50E VIB 194.00 0.23 30.00

777 APE 100 VIB 194.00 0.32 30.00

778 APE 100E VIB 194.00 0.14 30.00

779 APE 100HF VIB 260.00 0.14 43.00

780 APE 150 VIB 260.00 0.14 30.00

781 APE 150T VIB 260.00 0.17 30.00

782 APE 150HF VIB 466.00 0.32 43.00

783 APE 200 VIB 466.00 0.29 30.00

784 APE 200T VIB 466.00 0.34 30.83

785 APE 200T HF VIB 738.00 0.34 43.00

786 APE 300 VIB 738.00 0.34 25.00

787 APE 400B VIB 738.00 0.78 23.33

788 APE 600 VIB 800.00 1.05 23.30

789 APE Tan 400 VIB 1476.00 1.37 23.33

790 APE Tan 600 VIB 1800.00 2.11 23.30

791 APE 200-6 VIB 470.00 0.43 30.00

811 MGF RBH 80 VIB 50.00 0.60 30.00

812 MGF RBH 140 VIB 85.00 1.04 26.67

813 MGF RBH 200 VIB 125.00 0.74 26.67

814 MGF RBH 320 VIB 200.00 0.79 26.67

815 MGF RBH 460 VIB 255.00 1.13 26.67

816 MGF RBH 1050 VIB 460.00 1.55 22.50

817 MGF RBH 1575 VIB 700.00 1.16 22.50

818 MGF RBH 2400 VIB 975.00 1.77 23.50

880 ICE 23RF VIB 384.00 0.83 38.30

881 ICE 1412BT VIB 1193.00 1.67 21.70

882 ICE 23-40 VIB 30.00 0.19 31.80

883 ICE 28-35 VIB 261.00 1.16 27.30

884 ICE 28RF-35 VIB 261.00 1.16 27.30

885 ICE V360 VIB 783.00 0.94 25.00

886 ICE V360 T VIB 1566.00 1.88 25.00
539

Type Hammer 

Make 

Hammer 

Model

Hammer 

Type

Rated Power

kW 

Ecc. Mass

kips 

Frequency

Hz 

887 ICE 44-30V VIB 250.00 0.92 26.00

888 ICE 44-70 VIB 585.00 0.92 28.10

889 ICE 66-70 VIB 585.00 0.92 23.00

890 ICE 7RF VIB 154.00 0.51 38.30

891 ICE 66-70HS VIB 585.00 0.92 26.70

892 ICE 66-80HS VIB 597.00 0.92 29.20

893 ICE 100c-Tdm VIB 1774.00 1.83 26.67

894 ICE 423 VIB 377.00 0.92 38.30

895 ICE 32RF VIB 391.00 1.16 33.30

896 ICE 36RF VIB 431.00 1.30 33.30

897 ICE 46RF VIB 678.00 1.66 38.30

898 ICE 64RF VIB 663.00 1.16 32.50

899 ICE 44B VIB 595.00 1.30 30.00

900 Mueller MS16HF VIB 219.00 1.16 39.20

901 Mueller MS25H2 VIB 218.00 0.90 28.00

902 Mueller MS25H3 VIB 218.00 0.90 28.00

903 Mueller MS50H2 VIB 419.00 1.20 27.00

904 Mueller MS50H3 VIB 419.00 1.20 27.00

905 Mueller MS25HHF VIB 274.00 0.58 27.30

906 Mueller MS50HHF VIB 562.00 1.17 27.30

907 Mueller MS100HHF VIB 750.00 2.33 24.90

908 Mueller MS120HHF VIB 895.00 2.30 25.60

909 Mueller MS200HHF VIB 837.00 4.25 22.90

910 Mueller MS-10HFV VIB 203.00 0.39 39.30

911 Mueller MS-16HFV VIB 294.00 0.53 39.20

912 Mueller MS-24HFV VIB 720.00 0.85 39.20

913 Mueller MS-32HFV VIB 551.00 1.05 39.60

914 Mueller MS-48HFV VIB 823.00 1.69 39.20

915 Mueller MS-62HFV VIB 735.00 1.82 35.00

1039 J&M 11-23 VIB 164.00 0.92 31.70

1040 J&M 1412 VIB 559.00 1.67 21.70

1041 J&M 1412T VIB 1119.00 1.67 21.70

1042 J&M 216 VIB 149.00 0.92 26.70

1044 J&M 22-23 VIB 164.00 0.92 20.80

1045 J&M 22-30 VIB 261.00 0.92 27.50

1050 J&M 28-35 VIB 261.00 1.17 27.50

1051 J&M 360 VIB 783.00 0.94 21.70

1052 J&M 416 VIB 250.00 0.92 26.70

1053 J&M 416B VIB 261.00 0.92 26.70

1054 J&M 416S VIB 250.00 0.92 26.70

1055 J&M 815 VIB 429.00 0.92 26.70

1056 J&M 44-30 VIB 250.00 0.92 20.00

1057 J&M 44-50 VIB 399.00 0.92 26.70
540

Type Hammer 

Make 

Hammer 

Model

Hammer 

Type

Rated Power

kW 

Ecc. Mass

kips 

Frequency

Hz 

1058 J&M 44-65 VIB 552.00 0.92 27.50

1060 J&M 66-65 VIB 552.00 0.92 21.70

1061 J&M 66-80 VIB 559.00 0.92 26.70

1100 PVE 14M VIB 190.00 1.01 28.30

1101 PVE 23M VIB 234.00 1.66 27.50

1102 PVE 25M VIB 294.00 0.98 28.30

1103 PVE 27M VIB 294.00 0.98 28.30

1104 PVE 38M VIB 392.00 0.92 28.30

1105 PVE 50M VIB 440.00 1.20 28.30

1106 PVE 52M VIB 564.00 0.75 28.30

1107 PVE 105M VIB 784.00 1.52 22.50

1108 PVE 110M VIB 784.00 0.80 22.50

1109 PVE 200M VIB 1130.00 1.45 23.30

1110 PVE 2307 VIB 190.00 0.47 38.30

1111 PVE 1420 VIB 190.00 1.01 33.30

1112 PVE 2315 VIB 234.00 1.09 38.30

1113 PVE 2520 VIB 294.00 1.81 33.30

1114 PVE 2310VM VIB 190.00 0.72 38.30

1115 PVE 2315VM VIB 234.00 1.09 38.30

1116 PVE 2316VM VIB 294.00 1.16 38.30

1117 PVE 2319VM VIB 392.00 1.37 38.30

1118 PVE 2323VM VIB 392.00 0.83 38.30

1119 PVE 2332VM VIB 564.00 1.16 38.30

1120 PVE 2335VM VIB 784.00 1.27 38.30

1121 PVE 40VM VIB 564.00 1.45 33.30

1122 PVE 50VM VIB 564.00 1.20 30.00

1123 PVE 55M VIB 403.00 1.17 28.33

1124 PVE 82M VIB 565.00 1.76 28.33

1125 PVE 300M VIB 1796.00 6.21 23.33

1126 PVE 16VM VIB 335.00 0.35 38.33

1127 PVE 20VM VIB 395.00 0.41 38.33

1128 PVE 24VM VIB 395.00 0.52 38.33

1129 PVE 28VM VIB 403.00 0.61 38.33

1130 PVE 2070VM VIB 1130.00 1.52 33.33

1131 PVE 2312VM VIB 252.00 0.26 38.33

1132 PVE 2350VM VIB 790.00 1.09 38.33

1142 PTC 30HP VIB 196.00 0.87 27.00

1143 PTC 40HD VIB 269.00 0.87 28.00

1144 PTC 50HD1 VIB 255.00 0.87 25.00

1145 PTC 50HD2 VIB 290.00 0.87 25.00

1146 PTC 65HD VIB 305.00 0.87 26.00

1147 PTC 60HD VIB 305.00 0.87 28.00

1148 PTC 75HD VIB 410.00 0.87 25.00
541

Type Hammer 

Make 

Hammer 

Model

Hammer 

Type

Rated Power

kW 

Ecc. Mass

kips 

Frequency

Hz 

1149 PTC 100HD VIB 451.00 0.87 23.00

1150 PTC 100HDS VIB 564.00 0.87 23.00

1151 PTC 175HD VIB 611.00 0.87 23.00

1152 PTC 240HD VIB 988.00 0.87 23.00

1153 PTC 240HDS VIB 988.00 0.87 30.00

1154 PTC 120HD VIB 410.00 0.87 23.00

1155 PTC 130HD VIB 564.00 0.87 23.00

1156 PTC 200HD VIB 710.00 0.87 23.00

1157 PTC 265HD VIB 1080.00 0.87 24.00

1340 H&M H-150 VIB 94.00 0.11 28.30

1341 H&M H-1700 VIB 165.00 0.20 20.00

1431 BRUCE SGV-80 VIB 112.20 0.13 33.33

1432 BRUCE SGV-100 VIB 142.60 0.18 30.00

1433 BRUCE SGV-200 VIB 184.80 0.31 28.83

1434 BRUCE SGV-300 VIB 211.20 0.35 27.50

1435 BRUCE SGV-400 VIB 286.00 0.44 26.67

1436 BRUCE SGV-450 VIB 323.40 0.48 26.67

1437 BRUCE SGV-600 VIB 451.50 0.72 26.67

1438 BRUCE SGV-1000 VIB 569.10 1.03 25.00

1630 LBFoster 4150 VIB 335.00 0.53 25.00