Research, Development, and Technology

Turner-Fairbank Highway Research Center

6300 Georgetown Pike

McLean, VA 22101-2296

Postearthquake Reconnaissance Report 

on Transportation Infrastructure Impact 

of the February 27, 2010,

Offshore Maule Earthquake in Chile

Publication No. FHWA-HRT-11-030 October 2011
FOREWORD

On February 27, 2010, a devastating magnitude 8.8 earthquake struck off the coast of the Maule 

region of Chile, affecting a large area that included Chile’s two most populated cities: Concepción

and Santiago, the Chilean capital. A transportation infrastructure reconnaissance team (TIRT) 

was organized by the Federal Highway Administration (FHWA) and performed a thorough 

postearthquake investigation of highway infrastructure from April 4 to 13, 2010. TIRT was 

assisted by Chile’s Ministry of Public Works and two local universities: Pontifical Catholic 

University of Chile and University of Chile, both located in Santiago. This report presents the 

preliminary findings of the earthquake’s effects on the transportation infrastructure, including 

bridges and other highway structures that the team visited during the reconnaissance. This 

project was funded by FHWA’s Innovative Bridge Research and Deployment program. 

Jorge E. Pagán-Ortiz

Director, Office of Infrastructure

Research and Development

Notice

This document is disseminated under the sponsorship of the U.S. Department of Transportation 

in the interest of information exchange. The U.S. Government assumes no liability for the use 

of the information contained in this document. This report does not constitute a standard, 

specification, or regulation.

The U.S. Government does not endorse products or manufacturers. Trademarks or 

manufacturers’ names appear in this report only because they are considered essential to the 

objective of the document.

Quality Assurance Statement

The Federal Highway Administration (FHWA) provides high-quality information to serve 

Government, industry, and the public in a manner that promotes public understanding. Standards 

and policies are used to ensure and maximize the quality, objectivity, utility, and integrity of its 

information. FHWA periodically reviews quality issues and adjusts its programs and processes to 

ensure continuous quality improvement.
TECHNICAL DOCUMENTATION PAGE

1. Report No.

FHWA-HRT-11-030

2. Government Accession No. 3. Recipient’s Catalog No.

4. Title and Subtitle

Post-Earthquake Reconnaissance Report on Transportation Infrastructure: 

Impact of the February 27, 2010, Offshore Maule Earthquake in Chile

5. Report Date

October 2011

6. Performing Organization Code

7. Author(s)

Wen-Huei Phillip Yen, Genda Chen, Ian Buckle, Tony Allen, Daniel Alzamora, 

Jeffrey Ger, and Juan G. Arias

8. Performing Organization Report No.

9. Performing Organization Name and Address

Turner-Fairbank Highway Research Center

Federal Highway Administration

6300 Georgetown Pike, McLean, VA 22101

Department of Civil, Architectural and Environmental Engineering, 

Missouri University of Science and Technology

1401 N. Pine St., Rolla, MO 65409

Department of Civil and Environmental Engineering

University of Nevada Reno

1664 N. Virginia St., Reno, NV 89557

State Materials Laboratory

Washington State Department of Transportation

P.O. Box 47365, Olympia, WA 98504

10. Work Unit No. (TRAIS)

11. Contract or Grant No.

12. Sponsoring Agency Name and Address

Office of Infrastructure Research and Development

Federal Highway Administration

6300 Georgetown Pike

McLean, VA 22101-2296

13. Type of Report and Period Covered

Final Report: 

December 2009–Present

14. Sponsoring Agency Code

15. Supplementary Notes

The Contracting Officer’s Technical Representative (COTR) was W. Phillip Yen, HRDI-50.

16. Abstract

This report documents the findings and lessons learned from the February 27, 2010, M8.8 offshore Maule earthquake 

in Chile. Fewer than 0.15 percent of the bridges in Chile’s inventory, most built after 1995, collapsed or suffered 

damage that rendered them useless. Many spans of precast prestressed discontinuous girder bridges with continuous 

decks fell off their supports, probably due to significant in-plane rotation of the superstructure as a result of severe 

shaking. Lateral steel stoppers used to provide both vertical and lateral restraints on girders were largely unsuccessful 

due to their inadequate connection detail to cap beams and abutments. Reinforced concrete shear keys performed well 

as fuses limiting the transfer of excessive seismic loads from the superstructure to the foundation of bridges even 

though they could be optimized for maximum energy dissipation as part of the lateral restraint system at the bottom 

flange of girders. Vertical seismic bars were widely used to restrain the vertical motion of decks, and they also 

performed well. Bridge substructures (foundation, column, and cap beam) generally behaved satisfactorily except for 

two columns that suffered shear failure due to ground settlement and lateral spreading. All mechanically stabilized 

earth walls exceeded the expected performance.

17. Key Words

Seismic performance, Bridge damage, Superstructure rotation, Retaining 

wall, Soil liquefaction, Lateral spreading

18. Distribution Statement

No restrictions. This document is available 

to the public through NTIS:

National Technical Information Service

5301 Shawnee Road

Alexandria, VA 22312

19. Security Classif. (of this report)

Unclassified

20. Security Classif. (of this page)

Unclassified

21. No. of Pages

214

22. Price

Form DOT F 1700.7 (8-72) Reproduction of completed page authorized
ii

SI* (MODERN METRIC) CONVERSION FACTORS 

APPROXIMATE CONVERSIONS TO SI UNITS

Symbol When You Know Multiply By To Find Symbol 

LENGTH

in inches 25.4 millimeters mm 

ft feet 0.305 meters m 

yd yards 0.914 meters m 

mi miles 1.61 kilometers km 

AREA 

in2 square inches 645.2 square millimeters mm2

ft2 square feet 0.093 square meters m2

yd2 square yard 0.836 square meters m2

ac acres 0.405 hectares ha

mi2 square miles 2.59 square kilometers km2

VOLUME 

fl oz fluid ounces 29.57 milliliters mL 

gal gallons 3.785 liters L 

ft3 cubic feet 0.028 cubic meters m3 

yd3 cubic yards 0.765 cubic meters m3 

NOTE: volumes greater than 1000 L shall be shown in m3

MASS oz ounces 28.35 grams g

lb pounds 0.454 kilograms kg

T short tons (2000 lb) 0.907 megagrams (or "metric ton") Mg (or "t") 

TEMPERATURE (exact degrees) o

F Fahrenheit 5 (F-32)/9 Celsius o

C 

or (F-32)/1.8 

ILLUMINATION 

fc foot-candles 10.76 lux lx

fl foot-Lamberts 3.426 candela/m2 cd/m2

FORCE and PRESSURE or STRESS 

lbf poundforce 4.45 newtons N 

lbf/in2 poundforce per square inch 6.89 kilopascals kPa 

APPROXIMATE CONVERSIONS FROM SI UNITS 

Symbol When You Know Multiply By To Find Symbol 

LENGTH mm millimeters 0.039 inches in 

m meters 3.28 feet ft 

m meters 1.09 yards yd 

km kilometers 0.621 miles mi 

AREA mm2 square millimeters 0.0016 square inches in2 

m2 square meters 10.764 square feet ft2 

m2 square meters 1.195 square yards yd2 

ha hectares 2.47 acres ac

km2 square kilometers 0.386 square miles mi2 

VOLUME 

mL milliliters 0.034 fluid ounces fl oz 

L liters 0.264 gallons gal 

m3 cubic meters 35.314 cubic feet ft3 

m3 cubic meters 1.307 cubic yards yd3 

MASS 

g grams 0.035 ounces oz

kg kilograms 2.202 pounds lb

Mg (or "t") megagrams (or "metric ton") 1.103 short tons (2000 lb) T 

TEMPERATURE (exact degrees) o

C Celsius 1.8C+32 Fahrenheit oF 

ILLUMINATION 

lx lux 0.0929 foot-candles fc 

cd/m2 candela/m2 0.2919 foot-Lamberts fl

FORCE and PRESSURE or STRESS 

N newtons 0.225 poundforce lbf 

kPa kilopascals 0.145 poundforce per square inch lbf/in2

*SI is the symbol for the International System of Units. Appropriate rounding should be made to comply with Section 4 of ASTM E380. 

(Revised March 2003)
iii

TABLE OF CONTENTS

CHAPTER 1. INTRODUCTION.................................................................................................1

1.1 BACKGROUND ................................................................................................................1

1.2 THE TRANSPORTATION INFRASTRUCTURE RECONNAISSANCE TEAM ....4

1.2.1 Objective .....................................................................................................................4

1.2.2 Team Members ...........................................................................................................4

1.3 REPORT ORGANIZATION............................................................................................5

CHAPTER 2. GEOLOGY, SEISMOLOGY, AND GROUND MOTIONS .............................7

2.1 BACKGROUND ................................................................................................................7

2.2 GEOLOGY AND SEISMOLOGY...................................................................................7

2.3 GROUND MOTIONS .....................................................................................................10

CHAPTER 3. OVERVIEW OF BRIDGE PERFORMANCE AND SEISMIC DESIGN 

REQUIREMENTS IN CHILE ...................................................................................................23

3.1 OVERVIEW OF BRIDGE PERFORMANCE.............................................................23

3.1.1 Damage Statistics......................................................................................................23

3.1.2 Bridges Visited By Reconnaissance Team ...............................................................23

3.1.3 Damage Summary.....................................................................................................26

3.2. SEISMIC DESIGN REQUIREMENTS FOR BRIDGES...........................................27

CHAPTER 4. STRUCTURAL PERFORMANCE OF HIGHWAY BRIDGES....................29

4.1 OVERVIEW.....................................................................................................................29

4.2 PERFORMANCE OF STRAIGHT CONCRETE BRIDGES.....................................30

4.2.1 Américo Vespucio/Independencia ............................................................................30

4.2.2 Avenida Chada Acceso Sur Overpass.......................................................................34

4.2.3 Las Mercedes Route 5 Overpass...............................................................................35

4.2.4 Llacolen Bridge Over Biobío River..........................................................................36

4.2.5 Juan Pablo II Bridge Over Biobío River...................................................................41

4.2.6 Ramadillas Bridges...................................................................................................46

4.3 PERFORMANCE OF SKEWED AND CURVED CONCRETE AND STEEL 

BRIDGES................................................................................................................................51

4.3.1 Américo Vespucio/Miraflores ..................................................................................51

4.3.2 Américo Vespucio/Lo Echevers...............................................................................53

4.3.3 Avenida Romero Acceso Sur Overpass....................................................................55

4.3.4 Route 5 Railway Overcrossing at Hospital...............................................................57

4.3.5 Quilicura Railway Overcrossing at Avenida Manuel Antonio Matta.......................59

4.4 PERFORMANCE OF STEEL BRIDGES ....................................................................62

4.4.1 Tubul Bridge .............................................................................................................62

4.4.2 Cardenal Raúl Silva Henríquez Bridge.....................................................................65

4.4.3 Biobío River Bridge..................................................................................................71

4.4.4 Pichibudis Bridge......................................................................................................73

4.4.5 El Bar Bridge ............................................................................................................76

4.4.6 Itata River Bridge......................................................................................................79

4.4.7 Pedestrian Bridge Over Route 5 ...............................................................................81
iv

4.5 PERFORMANCE OF OTHER BRIDGE TYPES .......................................................82

4.5.1 Claro River Bridges..................................................................................................82

4.5.2 Chepe Railroad Bridge Over Biobío River...............................................................87

4.5.3 Maipú River Bridges.................................................................................................90

CHAPTER 5. GEOTECHNICAL PERFORMANCE OF TRANSPORTATION 

INFRASTRUCTURE ..................................................................................................................93

5.1 OVERVIEW.....................................................................................................................93

5.2 GEOTECHNICAL OBSERVATIONS REGARDING BRIDGE AND PIER 

STRUCTURES.......................................................................................................................94

5.2.1 Non-Liquefaction Geotechnical Issues Observed for Bridges..................................95

5.2.2 Liquefaction Impact on Performance of Bridge and Pier Structures........................99

5.2.3 Effect of Regional Ground Uplift Observed for Bridges........................................136

5.3 GEOTECHNICAL OBSERVATIONS REGARDING RETAINING WALL 

PERFORMANCE................................................................................................................138

CHAPTER 6. LESSONS LEARNED FROM THE EARTHQUAKE..................................155

6.1 GENERAL......................................................................................................................155

6.2 SUPERSTRUCTURE ROTATION.............................................................................155

6.2.1 Skewed Bridges ......................................................................................................155

6.2.2 Straight Bridges ......................................................................................................157

6.3 GIRDER DAMAGE ......................................................................................................158

6.3.1 Fracture of Steel Girders.........................................................................................158

6.3.2 Failure of Concrete Girders ....................................................................................158

6.4 CONNECTION BETWEEN SUPERSTRUCTURE AND SUBSTRUCTURE.......159

6.4.1 Shear Key and Steel Stopper Failures.....................................................................159

6.4.2 Vertical Seismic Bars..............................................................................................159

6.4.3 Bridge Bearings ......................................................................................................160

6.5 GIRDER SEAT LENGTH............................................................................................160

6.6 COLUMN SHEAR FAILURE......................................................................................160

6.7 FOUNDATION MOVEMENT AND DAMAGE........................................................160

6.7.1 Overview.................................................................................................................160

6.7.2 Sites Not Affected by Liquefaction ........................................................................160

6.7.3 Sites Affected by Liquefaction ...............................................................................161

6.8 RETAINING WALLS AND ROADWAY FILL ........................................................161

CHAPTER 7. CONCLUSIONS AND RECOMMENDATIONS ..........................................163

7.1 GENERAL......................................................................................................................163

7.2 CONCLUSIONS ON TRANSPORTATION INFRASTRUCTURE 

PERFORMANCE................................................................................................................163

7.3 RECOMMENDATIONS FOR IMMEDIATE IMPLEMENTATION.....................163

7.3.1 Diaphragms in Concrete Girder Superstructures....................................................163

7.3.2 Support Length........................................................................................................164

7.3.3 Lateral Seismic Restraints.......................................................................................164

7.3.4 Substructure Design................................................................................................164

7.3.5 Retaining Wall Design............................................................................................164

7.4 RECOMMENDATIONS FOR FUTURE WORK AND RESEARCH NEEDS ......165

7.4.1 Structural Engineering Recommendations .............................................................165

7.4.2 Geotechnical Engineering Recommendations........................................................167
v

APPENDIX A. SUBSURFACE DATA....................................................................................169

A.1 SUBSURFACE DATA..................................................................................................169

A.1.1 Subsurface Data for Américo Vespucio/Miraflores (Site 1)..................................169

A.1.2 Subsurface data for Américo Vespucio/Lo Echevers (Site 2) ...............................170

A.1.3 Subsurface Data for Avenida Romero Accesso Sur (Site 7) .................................171

A.1.4 Subsurface Data for Avenida Chada Accesso Sur (Site 8)....................................173

A.1.5 Subsurface Data for Route 5 Railway Overcrossing at Hospital (Site 10)............176

A.1.6 Subsurface Data for Via Elevada 21 de Mayo/Cruce Ferroviario (Site 28) ..........177

APPENDIX B. SEISMIC DESIGN REQUIREMENTS FOR BRIDGES AND WALLS 

IN CHILE...................................................................................................................................185

B.1 SEISMIC DESIGN REQUIREMENTS FOR BRIDGES .........................................185

B.1.1 Analysis Methods...................................................................................................186

B.1.2 Response Modification Factors (R)........................................................................189

B.1.3 Determination of Elastic Forces and Displacements..............................................190

B.1.4 Combination of Forces in Two Perpendicular Directions......................................190

B.1.5 Minimum Support Length......................................................................................191

B.1.6 Modified Design Forces for Structural Elements and Connections.......................191

B.1.7 Modified Design Forces for Foundations...............................................................191

B.1.8 Seismic Design of Foundations, Abutments, and Retaining Walls........................192

B.1.9 Seismic Design of Reinforced Concrete Elements ................................................192

B.1.10 Forces from Plastic Hinges on Piers or Columns.................................................192

B.1.11 Diaphragms ..........................................................................................................192

B.1.12 Anchor Bars (Seismic Bars).................................................................................193

B.1.13 Lateral Stoppers (Shear Keys)..............................................................................193

B.1.14 Seismic Joints.......................................................................................................193

B.1.15 Seismic Isolation Systems....................................................................................193

B.2. SEISMIC DESIGN REQUIREMENTS FOR WALLS............................................193

ACKNOWLEDGMENTS.........................................................................................................195

REFERENCES...........................................................................................................................197
vi

LIST OF FIGURES

Figure 1. Map. Site locations visited by TIRT.................................................................................2

Figure 2. Photo. TIRT team members and Chilean support personnel............................................5

Figure 3. Map. Location of epicenter, seismic recording stations, and TIRT sites.........................7

Figure 4. Map. Historic seismicity for Chile ...................................................................................8

Figure 5. Illustration. Conceptual diagram of a subduction zone earthquake .................................9

Figure 6. Illustration. Offshore Maule earthquake aftershock distribution with depth .................10

Figure 7. Illustration. Accelerogram from Universidad de Chile, Depto. Ing. Civil

(interior building)...........................................................................................................................12

Figure 8. Illustration. Response spectra from Universidad de Chile, Depto. Ing. Civil 

(interior building)...........................................................................................................................12

Figure 9. Illustration. Accelerogram from CRS Maipú RM..........................................................13

Figure 10. Illustration. Response spectra from CRS Maipú RM...................................................13

Figure 11. Illustration. Accelerogram from Hospital de Curicó....................................................14

Figure 12. Illustration. Response spectra from Hospital de Curicó...............................................14

Figure 13. Illustration. Accelerogram from San Pedro De La Paz, Colegio Concepción .............15

Figure 14. Illustration. Response spectra from San Pedro De La Paz, Colegio Concepción ........15

Figure 15. Map. Locations of ground motion sensors and structures visited by TIRT—

Santiago vicinity ............................................................................................................................16

Figure 16. Map. Locations of ground motion sensors and structure visited by TIRT—

Santiago to Rancagua.....................................................................................................................17

Figure 17. Map. Locations of ground motion sensors and structures visited by TIRT—

Curicó, Talca, Iloca vicinity...........................................................................................................18

Figure 18. Map. Locations of ground motion sensors and structures visited by TIRT—

Concepción vicinity .......................................................................................................................19

Figure 19. Map. Locations of ground motion sensors and structures visited by TIRT—

Maule epicenter to Tubul vicinity..................................................................................................20

Figure 20. Illustration. Geologic cross section of Concepción......................................................22

Figure 21. Map. Damaged sections in downtown Concepción .....................................................22

Figure 22. Map. Seismic zone map for central Chile ....................................................................28

Figure 23. Photo. Damage to eastbound Independencia bridge.....................................................30

Figure 24. Photo. Curtain wall at east abutment of eastbound Independencia bridge...................30

Figure 25. Photo. Damage to steel stoppers on eastbound Independencia bridge.........................31

Figure 26. Illustration. Details of steel stoppers............................................................................31

Figure 27. Photo. Typical flared wall pier of westbound Independencia bridge...........................32

Figure 28. Photo. Displaced seismic bar on westbound Independencia bridge.............................32

Figure 29. Photo. Shear key damage at abutment of westbound Independencia bridge ...............32

Figure 30. Photo. Shear key damage at intermediate bents of westbound Independencia bridge ...32

Figure 31. Photo. Westbound Independencia bridge .....................................................................33

Figure 32. Photo. Entrance ramp to westbound Independencia bridge .........................................33

Figure 33. Photo. Shear key damage at entrance ramp abutment..................................................33

Figure 34. Photo. Flared wall pier of exit ramp from westbound Independencia bridge ..............33

Figure 35. Photo. Close-up of damage to exit ramp from westbound Independencia bridge........34

Figure 36. Photo. Curtain wall shear failure at abutment of Chada bridge ...................................34

Figure 37. Photo. Bottom flange damage to Chada bridge............................................................34
vii

Figure 38. Photo. Curtain wall damage at bent of Chada bridge...................................................35

Figure 39. Photo. Soil separation from column of Chada bridge ..................................................35

Figure 40. Illustration. Plan view of superstructure rotation of Chada bridge ..............................35

Figure 41. Photo. Las Mercedes bridge and girder unseating at abutments..................................36

Figure 42. Photo. Llacolen bridge looking south-southwest .........................................................36

Figure 43. Photo. Unseated simply supported span in eastern approach to Llacolen bridge ........37

Figure 44. Photo. Abutment at unseated end of span in eastern approach to Llacolen bridge ......37

Figure 45. Photo. Concrete spalling beneath cap beam of eastern approach to Llacolen bridge.....37

Figure 46. Photo. Flexural cracks at the level of rip rap on eastern approach to Llacolen bridge ...38

Figure 47. Photo. Close-up view of cracks at the level of rip rap on eastern approach to 

Llacolen bridge ..............................................................................................................................38

Figure 48. Photo. Ground settlement and lateral movement of eastern approach to Llacolen 

bridge .............................................................................................................................................38

Figure 49. Photo. Exterior face of westbound ramp to Llacolen bridge at the river end...............39

Figure 50. Photo. Interior face of westbound ramp to Llacolen bridge at the river end................39

Figure 51. Photo. Bent of westbound ramp to Llacolen bridge .....................................................39

Figure 52. Photo. Seismic bar condition on westbound ramp to Llacolen bridge .........................39

Figure 53. Photo. Neoprene pad degradation on westbound ramp to Llacolen bridge..................40

Figure 54. Photo. Columns with soil marks indicating ground settlement at west end of 

Llacolen bridge ..............................................................................................................................40

Figure 55. Photo. Sand boils near columns at west end of Llacolen bridge ..................................40

Figure 56. Photo. Spalling in deck slab at west end of Llacolen bridge........................................40

Figure 57. Photo. Horizontal crack in web of end girders at west end of Llacolen bridge ...........41

Figure 58. Photo. Spalling of shear key at southwest abutment of Llacolen bridge......................41

Figure 59. Photo. Approaches to Juan Pablo II bridge ..................................................................41

Figure 60. Photo. Southwest view of northern approach to Juan Pablo II bridge, facing 

Biobío River...................................................................................................................................42

Figure 61. Photo. Northeast view of northern approach to Juan Pablo II bridge, facing away 

from Biobío River..........................................................................................................................42

Figure 62. Photo. Intermediate bent of northern approach to Juan Pablo II bridge.......................42

Figure 63. Photo. Column settlement under approach to Juan Pablo II bridge (cracks on far side)...43

Figure 64. Photo. Reduced column height due to shear failure under approach to 

Juan Pablo II bridge .......................................................................................................................43

Figure 65. Photo. Soil surface on far side of undamaged column under approach to 

Juan Pablo II bridge .......................................................................................................................43

Figure 66. Photo. Cracks on far side of undamaged column under approach to 

Juan Pablo II bridge .......................................................................................................................44

Figure 67. Photo. Differential settlement underneath first span over water at northern end of 

Juan Pablo II bridge .......................................................................................................................44

Figure 68. Photo. Shear failure in upstream column at northern end of Juan Pablo II bridge.......45

Figure 69. Photo. Front face of failure plane at northern end of Juan Pablo II bridge ..................45

Figure 70. Photo. Back face of failure plane at northern end of Juan Pablo II bridge...................45

Figure 71. Photo. Shear failure in downstream column at northern end of Juan Pablo II bridge ....46

Figure 72. Photo. Satellite image of Ramadillas bridges...............................................................46

Figure 73. Photo. Side view of east Ramadillas bridge .................................................................47

Figure 74. Photo. Wall-type pier beneath east Ramadillas bridge.................................................47

Figure 75. Photo. Deck misalignment under east Ramadillas bridge ............................................47
viii

Figure 76. Photo. Sand boils close to pier under east Ramadillas bridge......................................48

Figure 77. Photo. Overview of west Ramadillas bridge ................................................................48

Figure 78. Photo. Abutment damage beneath west Ramadillas bridge .........................................49

Figure 79. Photo. Pier rotation in collapsed span of west Ramadillas bridge ...............................49

Figure 80. Photo. Sand boils near collapsed span in west Ramadillas bridge ...............................49

Figure 81. Photo. Settlement at south abutment under east Ramadillas bridge.............................50

Figure 82. Photo. Sliding at south abutment under west Ramadillas bridge .................................50

Figure 83. Photo. Block flow toward river due to liquefaction on south bank of Ramadillas 

bridges............................................................................................................................................50

Figure 84. Photo. Northeast end of Miraflores bridge ...................................................................51

Figure 85. Photo. Southwest end of Miraflores bridge..................................................................51

Figure 86. Photo. Top view of Miraflores bridge from far side ....................................................51

Figure 87. Photo. Bottom view of Miraflores bridge from near side ............................................52

Figure 88. Photo. Collapsed span of Miraflores bridge .................................................................52

Figure 89. Photo. Miraflores bridge after removal of superstructure ............................................52

Figure 90. Photo. Acute corner of southwest abutment of Miraflores bridge ...............................53

Figure 91. Photo. Obtuse corner of southwest abutment of Miraflores bridge..............................53

Figure 92. Photo. Collapsed northeast-bound Lo Echevers bridge ...............................................53

Figure 93. Photo. Three unseated spans of northeast-bound Lo Echevers bridge.........................54

Figure 94. Photo. Failure of a steel stopper on Lo Echevers bridge..............................................54

Figure 95. Photo. Displaced elastomeric bearing on Lo Echevers bridge .....................................54

Figure 96. Illustration. Intermediate bent details for Romero bridge ............................................55

Figure 97. Photo. Collapse of Romero bridge ...............................................................................55

Figure 98. Photo. Seismic bars at west abutment of Romero bridge.............................................56

Figure 99. Photo. Seismic bars at east abutment of Romero bridge ..............................................56

Figure 100. Photo. Pounding at west abutment of Romero bridge................................................56

Figure 101. Photo. Back wall and wing wall damage at west abutment of Romero bridge ..........57

Figure 102. Photo. Three bridges (one demolished) at Route 5 overcrossing at Hospital ............57

Figure 103. Photo. Collapse of bridge at Route 5 overcrossing at Hospital..................................57

Figure 104. Photo. Damage at north abutment of Route 5 overcrossing at Hospital ....................58

Figure 105. Photo. Damage at south abutment of Route 5 overcrossing at Hospital ....................59

Figure 106. Photo. Overview of Quilicura railway overcrossing ..................................................59

Figure 107. Photo. Cross frames on Quilicura railway overcrossing ............................................60

Figure 108. Photo. Stiffeners on Quilicura railway overcrossing..................................................60

Figure 109. Photo. East abutment of Quilicura railway overcrossing ...........................................60

Figure 110. Photo. West abutment of Quilicura railway overcrossing..........................................61

Figure 111. Photo. West span of Quilicura railway overcrossing lowered onto sandbank ...........61

Figure 112. Photo. Damage to anchor bolts on Quilicura railway overcrossing ...........................61

Figure 113. Photo. Damage to cross frame on Quilicura railway overcrossing ............................62

Figure 114. Photo. Unseating at southern ends of each span of Tubul bridge ..............................62

Figure 115. Photo. Punching of Tubul bridge deck into back wall ...............................................62

Figure 116. Photo. Buckling of pavement at north abutment of Tubul bridge..............................63

Figure 117. Photo. Span 6 of Tubul bridge, unseated at north end during an aftershock..............63

Figure 118. Photo. Bottom flange buckling and concrete spalling on Tubul bridge .....................64

Figure 119. Photo. Shear crack in Tubul bridge ............................................................................64

Figure 120. Photo. Tilting wall pier on Tubul bridge ....................................................................64

Figure 121. Photo. Spalling due to shear action and vertical crack in Tubul bridge .....................65
ix

Figure 122. Photo. Crack in footing of Tubul bridge ....................................................................65

Figure 123. Photo. Cardenal Raúl Silva Henríquez bridge looking south.....................................66

Figure 124. Photo. Elastomeric pad and stopper over bent on Cardenal Raúl Silva Henríquez 

bridge .............................................................................................................................................66

Figure 125. Photo. Girder offset and cross frame buckling on Cardenal Raúl Silva Henríquez 

bridge .............................................................................................................................................67

Figure 126. Photo. Cross frame buckling on Cardenal Raúl Silva Henríquez bridge ...................67

Figure 127. Photo. Girder damage at northeast abutment of Cardenal Raúl Silva Henríquez 

bridge .............................................................................................................................................68

Figure 128. Photo. Temporary repair at northeast abutment of Cardenal Raúl Silva Henríquez

bridge .............................................................................................................................................68

Figure 129. Photo. Spans and pile bents of southwest portion of Cardenal Raúl Silva Henríquez

bridge .............................................................................................................................................69

Figure 130. Photo. Erosion of alluvial material around bent legs of Cardenal Raúl Silva 

Henríquez bridge............................................................................................................................69

Figure 131. Photo. Debris-impact hole in bent leg of Cardenal Raúl Silva Henríquez bridge......69

Figure 132. Photo. Fillet weld fractures at the girder-to-abutment connection on the Cardenal 

Raúl Silva Henríquez bridge..........................................................................................................70

Figure 133. Photo. Close-up of girder end on Cardenal Raúl Silva Henríquez bridge..................70

Figure 134. Photo. Close-up of transverse stiffeners on Cardenal Raúl Silva Henríquez bridge ....70

Figure 135. Photo. Collapsed spans in Biobío River bridge..........................................................71

Figure 136. Photo. Unseated spans in Biobío River bridge...........................................................72

Figure 137. Photo. Lateral spreading near east end of Biobío River bridge..................................72

Figure 138. Photo. Collapsed spans and piers of Biobío River bridge ..........................................72

Figure 139. Photo. Cracks in pile caps of Biobío River bridge .....................................................72

Figure 140. Photo. Satellite image of Pichibudis bridge ...............................................................73

Figure 141. Photo. Cross frames in Pichibudis bridge...................................................................73

Figure 142. Photo. Web stiffeners in Pichibudis bridge ................................................................74

Figure 143. Photo. Offset in girder top flange of south abutment of Pichibudis bridge................74

Figure 144. Photo. Offset in handrail on south abutment of Pichibudis bridge.............................74

Figure 145. Photo. X-braces damage in south abutment of Pichibudis bridge..............................75

Figure 146. Photo. Diaphragm crushing in south abutment of Pichibudis bridge.........................75

Figure 147. Photo. Damage on south abutment due to corrosion in Pichibudis bridge.................76

Figure 148. Photo. Satellite image of El Bar bridge ......................................................................76

Figure 149. Photo. Side view of El Bar bridge ..............................................................................77

Figure 150. Photo. Retrofit girder connections on El Bar bridge ..................................................77

Figure 151. Photo. Retrofit deck connections on El Bar bridge ....................................................77

Figure 152. Photo. Lateral deformation at south abutment of El Bar bridge.................................78

Figure 153. Photo. Damage at south abutment of El Bar bridge ...................................................78

Figure 154. Photo. Damage to rebar X-brace system on north abutment of El Bar bridge...........78

Figure 155. Photo. Damage to north abutment back wall of El Bar bridge...................................79

Figure 156. Photo. Satellite image of Itata River bridge ...............................................................79

Figure 157. Photo. Side view and superstructure details for Itata River bridge ............................80

Figure 158. Photo. Girder supports at each pier of Itata River bridge...........................................80

Figure 159. Photo. Damage to girder supports of Itata River bridge.............................................81

Figure 160. Photo. Damaged Route 5 pedestrian bridge ...............................................................81

Figure 161. Photo. Steel girder on neoprene pad at abutment of Route 5 pedestrian bridge ........82
x

Figure 162. Photo. End cross frame near collapsed span of Route 5 pedestrian bridge................82

Figure 163. Photo. Anchor bolt failure (shear and pull out) on Route 5 pedestrian bridge...........82

Figure 164. Illustration. Highway bridges crossing Claro River before earthquake .....................83

Figure 165. Photo. Collapse of masonry arch bridge at Claro River.............................................83

Figure 166. Photo. Close-up view of collapsed bridge at Claro River ..........................................83

Figure 167. Photo. Deck slab and crown of arch on Claro River bridge.......................................84

Figure 168. Photo. Pier of Claro River bridge ...............................................................................84

Figure 169. Photo. Foundation of Claro River bridge ...................................................................85

Figure 170. Illustration. Scenario for the progressive collapse of Claro River bridge ..................86

Figure 171. Photo. North end of Chepe railroad bridge over roadway..........................................87

Figure 172. Photo. River spans of Chepe railroad bridge..............................................................87

Figure 173. Photo. South end of Chepe railroad bridge ................................................................87

Figure 174. Photo. Bottom view and cross section on north end of Chepe railroad bridge ..........87

Figure 175. Photo. Section on south end of Chepe railroad bridge ...............................................88

Figure 176. Illustration. Settlement and lateral movement of wall and bent due to ground 

spreading under Chepe railroad bridge ..........................................................................................88

Figure 177. Photo. Settlement and tilting of bent on Chepe railroad bridge .................................89

Figure 178. Photo. Settlement behind retaining wall on Chepe railroad bridge............................89

Figure 179. Photo. Ruptured cross bracing at south end of Chepe railroad bridge .......................90

Figure 180. Photo. Fractured steel piles at south end of Chepe railroad bridge ............................90

Figure 181. Photo. Three parallel bridges crossing Maipú River south of Santiago.....................90

Figure 182. Photo. Crushing of haunches at pier 3 of Maipú River bridge...................................91

Figure 183. Photo. Shear cracks at span 1 of Maipú River bridge ................................................91

Figure 184. Photo. Longitudinal and vertical seismic forces in haunches at pier 3 of 

Maipú River bridge ........................................................................................................................91

Figure 185. Photo. Damage at Romero bridge and shaft foundation supporting abutment 

(collapsed superstructure removed)...............................................................................................96

Figure 186. Photo. Interior pier of Chada structure with 5.8- to 7.8-inch (150- to 200-mm) 

gap in soil transverse to bridge centerline......................................................................................97

Figure 187. Photo. East abutment of Chada structure showing large pavement crack directly 

below bridge centerline..................................................................................................................97

Figure 188. Illustration. Fill slumping observed for bridge sites 7 and 8......................................98

Figure 189. Photo. Fill slumping observed at Romero bridge.......................................................98

Figure 190. Photo. View of roadway fill failure from abutment of Route 5 railway crossing 

at Hospital ......................................................................................................................................99

Figure 191. Photo. Close-up of approach fill failure and damaged sign bridge ............................99

Figure 192. Photo. Lateral spreading at Mataquito bridge ..........................................................101

Figure 193. Photo. Lateral spreading at Mataquito bridge with lateral movement beneath 

and around bridge ........................................................................................................................101

Figure 194. Photo. Lateral spreading and ground failure at old Ramadillas bridge ....................102

Figure 195. Photo. Sand boils at southwest end of Llacolen bridge............................................102

Figure 196. Photo. Ejected sand due to liquefaction and ground failure at La Mochita bridge ..103

Figure 197. Photo. Severe ground failure at La Mochita bridge .................................................104

Figure 198. Photo. Lateral spreading and ground failure at Raqui 2 bridge................................104

Figure 199. Photo. San Nicolás bridge center pier settlement probably due to liquefaction 

of soil layers beneath river...........................................................................................................105
xi

Figure 200. Photo. Ramps at Juan Pablo II bridge where existing ground was improved 

with DDC.....................................................................................................................................106

Figure 201. Photo. Lateral movement of fill in ramps at Juan Pablo II bridge ...........................106

Figure 202. Photo. Abutment fill settlement of northeast abutment of Mataquito bridge...........108

Figure 203. Photo. Soil settlement without bridge settlement at northeast abutment of 

Mataquito bridge..........................................................................................................................108

Figure 204. Photo. Soil settlement at interior pier at northeast abutment of Mataquito bridge...109

Figure 205. Photo. Abutment fill settlement at northwest abutment of Raqui 1 bridge ..............110

Figure 206. Photo. Abutment and girders shoved into each other and deck at the northwest 

abutment of the Raqui 1 bridge....................................................................................................111

Figure 207. Photo. Raqui 1 bridge viewed from north side.........................................................111

Figure 208. Photo. Evidence of ground failure perpendicular to roadway centerline at 

northwest approach of Raqui 1 bridge .........................................................................................112

Figure 209. Map. Aerial view of Biobío River and crossings.....................................................113

Figure 210. Photo. Evidence of slope instability due to liquefaction at northeast abutment 

and approaches of Llacolen bridge ..............................................................................................113

Figure 211. Illustration. Plan view of liquefaction-induced slide at northeast abutment of 

Llacolen bridge ............................................................................................................................114

Figure 212. Photo. Settlement due to liquefaction-induced slope instability at northeast end 

of Llacolen bridge ........................................................................................................................114

Figure 213. Photo. Settlement and lateral ground movement of interior pier at northeast end 

of Llacolen bridge ........................................................................................................................115

Figure 214. Photo. Lateral ground and foundation movement at interior pier and fallen span 

at northeast end of Llacolen bridge..............................................................................................116

Figure 215. Photo. Close-up of cracked column at northeast end of Llacolen bridge.................116

Figure 216. Photo. Wall settlement due to slope movement and instability caused by 

liquefaction at northwest end of Chepe railroad bridge...............................................................117

Figure 217. Photo. Settlement of approach structure for Juan Pablo II bridge............................118

Figure 218. Photo. Close-up of settlement of approach structure for Juan Pablo II bridge.........118

Figure 219. Photo. Roadway settlement due to settlement of approach structure pier 

foundations for Juan Pablo II bridge............................................................................................119

Figure 220. Photo. Roadway settlement due to settlement of mainline pier foundations for 

Juan Pablo II Bridge.....................................................................................................................119

Figure 221. Photo. Liquefaction-induced differential settlement of mainline pier foundations

supporting Juan Pablo II bridge ...................................................................................................120

Figure 222. Photo. Damage to pier columns for mainline Juan Pablo II bridge..........................120

Figure 223. Photo. View of collapsed spans from northeast river bank of Biobío River............121

Figure 224. Photo. Close-up of old Biobío River bridge showing exposed pile foundations .....121

Figure 225. Photo. Ground failure near northeast abutment of old Biobío River bridge ............122

Figure 226. Photo. Tilting of northeast abutment pier wall of old Biobío River bridge .............122

Figure 227. Photo. Ground failure of north approach fill for La Mochita bridge........................123

Figure 228. Photo. La Mochita bridge north approach fill liquefaction-induced ground failure...124

Figure 229. Photo. La Mochita bridge north approach fill and edge of dredging pit ..................124

Figure 230. Photo. La Mochita bridge deck and girder lateral movement ..................................125

Figure 231. Photo. La Mochita bridge deck and girder lateral movement and approach fill 

settlement .....................................................................................................................................126

Figure 232. Photo. La Mochita bridge lateral bridge deck and girder at south abutment ...........126
xii

Figure 233. Photo. Southeast approach fill damage at Raqui 2 bridge........................................127

Figure 234. Photo. Collapsed and shifted spans at Raqui 2 bridge as viewed from base of 

southeast abutment.......................................................................................................................128

Figure 235. Photo. Collapsed and shifted spans at Raqui 2 bridge due to liquefaction...............128

Figure 236. Photo. South approach fill damage and collapsed spans at Tubul Bridge................129

Figure 237. Photo. Settlement of fill beside the south abutment of Tubul bridge.......................129

Figure 238. Photo. Overview of damage to south abutment of Tubul bridge .............................130

Figure 239. Photo. Close-up of exposed pile and pile cap damage on south abutment of 

Tubul bridge.................................................................................................................................130

Figure 240. Photo. Overview of Tubul bridge north abutment....................................................131

Figure 241. Photo. Tubul bridge collapsed spans and substructure.............................................131

Figure 242. Photo. Puerto de Coronel Muelle foundation displacement and bending ................132

Figure 243. Photo. Close-up of Puerto de Coronel Muelle foundation displacement and 

bending.........................................................................................................................................132

Figure 244. Photo. View northeast from Tubul bridge south abutment showing areas left dry

and boats left above the new mean water level............................................................................137

Figure 245. Photo. Tubul bridge with boats run through by old timber pilings..........................137

Figure 246. Photo. Retaining wall at Américo Vespucio/Miraflores ..........................................139

Figure 247. Photo. Retaining wall at Américo Vespucio/Lo Echevers.......................................140

Figure 248. Photo. Retaining wall at 14 de la Fama....................................................................140

Figure 249. Photo. Retaining wall at Quilicura railroad overcrossing ........................................140

Figure 250. Photo. Retaining wall at Américo Vespucio/Independencia eastbound...................140

Figure 251. Photo. Retaining wall at Américo Vespucio/Independencia westbound .................140

Figure 252. Photo. Retaining wall at Maipú River bridge...........................................................141

Figure 253. Photo. Retaining wall at Estribo Francisco Mostazal (Avenida Independencia).....141

Figure 254. Photo. Retaining wall at Chepe railroad bridge over Biobío River..........................141

Figure 255. Photo. Retaining wall at Raqui 2..............................................................................141

Figure 256. Photo. South retaining wall at Juan Pablo II ............................................................141

Figure 257. Photo. North retaining wall at Juan Pablo II ............................................................142

Figure 258. Photo. Retaining wall at 21 de Mayo .......................................................................142

Figure 259. Photo. Retaining wall at Rotonda General Bonilla ..................................................142

Figure 260. Photo. Retaining wall at Muros Talca (SW) ............................................................142

Figure 261. Photo. Toppling of coping at Américo Vespucio/Independencia ............................143

Figure 262. Photo. Wall movement near Chepe railroad bridge due to lateral spreading and 

settlement .....................................................................................................................................144

Figure 263. Photo. Wall settlement and rotation due to lateral spreading and settlement...........144

Figure 264. Map. Retaining walls near Juan Pablo II bridge.......................................................145

Figure 265. Photo. Wall settlement near Juan Pablo II bridge due to liquefaction-induced 

settlement .....................................................................................................................................145

Figure 266. Map. Vía Elevada 21 de Mayo.................................................................................146

Figure 267. Photo. Tiered wall corner tilting outward at wall site 28A ......................................147

Figure 268. Photo. Wall corner tilting outward at wall site 28B.................................................148

Figure 269. Photo. Lateral movement of lower wall top at wall site 28C...................................149

Figure 270. Photo. Back side of upper wall at wall site 28C.......................................................149

Figure 271. Photo. Fill material from tiered wall at wall site 28C ..............................................150

Figure 272. Photo. Differential settlement at wall site 28C.........................................................150

Figure 273. Photo. Possible liquefaction features near wall site 28C..........................................151
xiii

Figure 274. Illustration. Plan view of wall at Muros Talca .........................................................152

Figure 275. Photo. Wall with steep top slope at Muros Talca.....................................................152

Figure 276. Photo. Wall with cracked block pattern at Muros Talca ..........................................153

Figure 277. Photo. Bridge supported on footing on top of block-faced geogrid wall at 

Estribo Francisco Mostazal..........................................................................................................153

Figure 278. Illustration. Notations on skew direction..................................................................155

Figure 279. Illustration. Deck rotation of a representative bridge (two spans shown)................156

Figure 280. Illustration. Rotational and translational mode responses to ground motions..........157

Figure 281. Photo. Exterior girder damage at Chada bridge, no diaphragms..............................158

Figure 282. Photo. Exterior girder damage at San Nicolás bridge, partial diaphragms ..............159

Figure 283. Photo. Exterior girder damage at west abutment of Llacolen bridge, concrete teeth...159

Figure 284. Map. Image of site 1 taken prior to earthquake........................................................169

Figure 285. Illustration. Boring log at site 1................................................................................169

Figure 286. Map. Image of site 2 taken prior to earthquake and showing approximate 

location of test hole......................................................................................................................170

Figure 287. Illustration. As-built bridge plan for site 2 ...............................................................170

Figure 288. Illustration. As-built bridge profile for site 2 showing footing foundations ............171

Figure 289. Illustration. Boring log at site 2................................................................................171

Figure 290. Map. Image of site 7 taken prior to earthquake and showing approximate 

location of test hole......................................................................................................................172

Figure 291. Illustration. As-built bridge plan for site 7 ...............................................................172

Figure 292. Illustration. Boring log for site 7 ..............................................................................173

Figure 293 Map. Image of site 8 taken prior to earthquake and showing approximate 

location of test hole......................................................................................................................173

Figure 294. Illustration. As-built bridge plan for site 8 ...............................................................174

Figure 295. Illustration. Boring log for site 8 ..............................................................................175

Figure 296. Map. Image of site 10 taken prior to earthquake and showing approximate 

location of test hole......................................................................................................................176

Figure 297. Illustration. As-built bridge plan for site 10 .............................................................176

Figure 298. Illustration. Boring log SA-2 for site 10...................................................................177

Figure 299. Map. Image of site 28 taken prior to earthquake and showing approximate 

locations of test holes...................................................................................................................177

Figure 300. Illustration. Boring log S-1 for site 28......................................................................178

Figure 301. Illustration. Detailed standard penetration test blow counts for boring log S-1 

for site 28 .....................................................................................................................................179

Figure 302. Illustration. Boring log S-2 for site 28......................................................................180

Figure 303. Illustration. Detailed standard penetration test blow counts for boring log S-2 

for site 28 .....................................................................................................................................181

Figure 304. Illustration. Boring log S-3 for site 28......................................................................182

Figure 305. Illustration. Detailed standard penetration test blow counts for boring log S-3 

for site 28 .....................................................................................................................................183
xiv

LIST OF TABLES

Table 1. Specific sites’ structural names and GPS locations...........................................................3

Table 2. Summary of available earthquake records for the offshore Maule earthquake ...............11

Table 3. Number of damaged bridges............................................................................................23

Table 4. Bridges visited by reconnaissance team ..........................................................................24

Table 5. Summary of performance for abutments and interior piers on slopes affected by 

liquefaction-induced ground movement or settlement ...............................................................134

Table 6. Summary of sites with retaining walls...........................................................................139

Table 7. Summary of bridges and bridge damage .......................................................................155

Table 8. Seismic zones.................................................................................................................185

Table 9. Soil coefficients .............................................................................................................185

Table 10. Seismic behavior categories.........................................................................................186

Table 11. Importance factor, K1 ...................................................................................................187

Table 12. Spectral constants, T1 and K2 .......................................................................................187

Table 13. Approximate formulas for natural periods for bridges................................................188

Table 14. Response modification factors (R)...............................................................................190
xv

LIST OF ACRONYMS

AASHTO American Association of State and Highway Transportation Officials

CIP Cast-in-place

DDC Deep dynamic compaction

EERI Earthquake Engineering Research Institute

FHWA Federal Highway Administration

GEER Geotechnical Extreme Events Reconnaissance

GPS Global Positioning System

LRFD Load and resistance factor design

MOP Ministry of Public Works

MSE Mechanically stabilized earth

NEHRP National Earthquake Hazards Reduction Program

PGA Peak ground acceleration

QDR Quake data recorder

TIRT Transportation infrastructure reconnaissance team

USGS United States Geological Survey

1

CHAPTER 1. INTRODUCTION

1.1 BACKGROUND

Over the course of the last century, Chile has experienced several strong earthquakes. According 

to the United States Geological Survey (USGS), M8.2 earthquakes struck the nation in 1906, 1943,

and 1960, and an M8.0 earthquake hit in 1985. The M8.2 earthquake in 1960 was a foreshock that

occurred the day before the largest earthquake ever recorded, the M9.5 Chilean earthquake. The

offshore Maule earthquake on February 27, 2010, was measured at M8.8 and lasted more than 

2 min. It is estimated that this earthquake was approximately 500 times more powerful than the 

earthquake that devastated Haiti in January 2010. Several aftershocks, including nine events with 

a magnitude exceeding 6.0, occurred in the days following the Chilean earthquake.

(1) The M8.8 

event is the fifth largest earthquake recorded in modern times and was characterized by its long 

duration and strong ground motion, which also caused tsunamis across the region. Many bridges 

and tunnels constructed with seismic design codes similar to the 1983 U.S. and European codes 

were damaged in the earthquake.

Because of the size of the event, the intensity of the ground shaking, the geological similarity of 

the Maule region and the Northwestern United States (e.g., Washington and Oregon’s subduction 

zones), and the comparable nature of the infrastructure construction and seismic design codes 

used, the performance of transportation infrastructure during this earthquake provided valuable

lessons in earthquake engineering. In March 2010, the Federal Highway Administration (FHWA)

contacted Chile’s Ministry of Public Works (MOP) and worked with University Transportation 

Center, Washington State Department of Transportation, University of Nevada, and Missouri 

University of Science and Technology to organize the transportation infrastructure reconnaissance

team (TIRT). TIRT and a member of the Earthquake Engineering Research Institute (EERI) were

dispatched to Chile on April 3 to perform a comprehensive earthquake reconnaissance. TIRT was

supported by local bridge engineers from MOP and local researchers from Pontifical Catholic 

University of Chile and the University of Chile.

TIRT visited more than 32 transportation infrastructure sites, including highway bridges and port

facilities from Santiago down to Tubul, a small city near Concepción. The specific locations visited

are shown in figure 1 and listed in table 1 with site names, Global Positioning System (GPS)

coordinates, and city and county names.
2

© Google, Inav/Geosistemas SRL, Europa Technologies, and DMapas

Figure 1. Map. Site locations visited by TIRT.
3

Table 1. Specific sites’ structural names and GPS locations.

Site 

No. Site Name Location

Latitude

(S) 

(degrees)

Longitude

(W) 

(degrees)

1a Américo Vespucio/Miraflores eastbound

1b Américo Vespucio/Miraflores westbound Santiago -33.39 -70.77

2a Américo Vespucio/Lo Echevers eastbound

2b Américo Vespucio/Lo Echevers westbound Santiago -33.38 -70.75

3 I-5/14 de la Fama Santiago -33.40 -70.68

4 Pedestrian bridge over Route 5 Santiago -33.34 -70.71

5

Quilicura railway crossing at Avenida 

Manuel Antonio Matta Santiago -33.37 -70.70

6a

Américo Vespucio/Independencia 

westbound

Santiago -33.37 -70.69

6b

Américo Vespucio/Independencia 

eastbound

6c

Américo Vespucio/Independencia 

westbound exit ramp

6d

Américo Vespucio/Independencia 

westbound entrance ramp

7 Avenida Romero Acceso Sur overpass Paine -33.86 -70.72

8 Avenida Chada Acceso Sur overpass Paine -33.87 -70.73

9a Maipú River (local access bridge)

Buin -33.69 -70.72

9b Maipú River (current Route 5 bridge)

9c Maipú railroad crossing

10a

Route 5 railway crossing at Hospital 

westbound

10b Buin -33.86 -70.75

Route 5 railway crossing at Hospital 

eastbound

11

Estribo Francisco Mostazal (Avenida

Independencia) Mostazal -34.03 -70.72

12 Las Mercedes Route 5 overpass Rancagua -34.07 -70.76

13a Claro River

13b Claro River San Rafael -35.18 -71.39

14 Pichibudis Iloca -34.88 -72.16

15 Mataquito Iloca -35.05 -72.16

16 Cardenal Raúl Silva Henríquez Constitución -35.34 -72.39

17 Llacolen Concepción -36.83 -73.07

18 Chepe railroad bridge over Biobío River Concepción -36.82 -73.07

19a Puerto de Coronel Muelle Norte

19b Puerto de Coronel Muelle Sur Coronel -37.03 -73.15

20 Raqui 1 Raqui -37.25 -73.44

21 Raqui 2 Raqui -37.25 -73.44
4

Table 1. Specific sites’ structural names and GPS locations—Continued.

Site 

No. Site Name Location

Latitude

(S) 

(degrees)

Longitude

(W) 

(degrees)

22 Tubul Raqui -37.23 -73.46

23 El Bar Arauco -37.26 -73.24

24a Ramadillas (west, old)

24b Ramadillas (east, new) Arauco -37.31 -73.26

25 Juan Pablo II Concepción -36.82 -73.09

26 Biobío River (old) Concepción -36.84 -73.06

27 La Mochita Concepción -36.85 -73.06

28 Vía Elevada 21 de Mayo/Cruce Ferroviario Concepción -36.82 -73.07

29 Rotonda General Bonilla Concepción -36.81 -73.03

30 Itata River Coelemu -36.47 -72.69

31 San Nicolás San Nicolás -36.50 -72.21

32 Muros Talca (SW) Talca -35.48 -71.67

1.2 THE TRANSPORTATION INFRASTRUCTURE RECONNAISSANCE TEAM

1.2.1 Objective

TIRT’s mission was to conduct a thorough postearthquake investigation concentrating on highway

bridges, tunnels, and retaining walls in the areas affected by the earthquake, including the cities 

of Concepción and Santiago. Performance of these structures, including damaged and undamaged

conditions, were carefully documented and analyzed. The information gathered from this effort 

will be studied, and the results will be used to assess, refine, and improve current design codes 

and standards that benefit the United States, Chile, and the general engineering community.

1.2.2 Team Members

TIRT included six members led by the FHWA Office of Infrastructure Research and Development,

including three FHWA representatives, one representative of the American Association of State 

and Highway Transportation Officials (AASHTO), and two university representatives. Juan G. 

Arias, a student representing EERI, joined and assisted the team in the investigation. The team

was coordinated by FHWA’s Sheila Duwadi throughout the preparation, reconnaissance, and 

postreconnaissance reporting. The U.S. team members were fully supported by their Chilean

colleagues. Figure 2 shows the team members and local participants. Pictured from left to right are

Daniel Alzamora (FHWA Resource Center), Dr. Ian Buckle (University of Nevada), Dr. Phillip 

W. Yen (FHWA team leader), Genda Chen (Missouri University of Science and Technology), 

Juan G. Arias (University of Nevada), Rodrigo Oviedo (Pontifical Catholic University of Chile),

Dr. Jeffrey Ger (FHWA Florida Division), Tony Allen (Washington State Department of 

Transportation), Sandra Achurra (MOP), and Mauricio Guzman (MOP).
5

Figure 2. Photo. TIRT team members and Chilean support personnel.

1.3 REPORT ORGANIZATION

This report presents the preliminary reconnaissance findings of the earthquake performance of 

transportation infrastructure based on available information, visual observations, and preliminary 

analysis. Chapter 2 describes the seismic details of the earthquake. Chapter 3 presents a summary 

of Chilean seismic design codes studied by TIRT, and chapter 4 provides the findings of the 

investigation of structural performance. Chapter 5 documents the geotechnical observations of 

infrastructure performance, including bridges, retaining walls, and other structures. Chapter 6 

presents the lessons learned from this reconnaissance, and chapter 7 offers conclusions and

recommendations.

7

CHAPTER 2. GEOLOGY, SEISMOLOGY, AND GROUND MOTIONS

2.1 BACKGROUND

The USGS Web site and a report provided by the Geotechnical Extreme Events Reconnaissance 

(GEER) team were used to provide necessary background information regarding the seismic and 

geologic setting for the sites specifically investigated by TIRT.(1,2) Additional information not 

available at the time of the USGS posting and the GEER report (e.g., boring logs, ground motion 

data, etc.) is also included in this chapter.

2.2 GEOLOGY AND SEISMOLOGY

The offshore Maule earthquake occurred on February 27, 2010, at 3:34 a.m. (local time) at the 

boundary between the Nazca and South American tectonic plates.(1) The location of the earthquake’s

epicenter is shown in figure 3 relative to the sites investigated by TIRT, shown as white dots with 

black centers and white numbers. The earthquake epicenter, located at 35.909 °S, 72.733 °W, was 

approximately 208 mi (335 km) southwest of Santiago, 65.2 mi (105 km) northeast of Concepción,

and 71.4 mi (115 km) west-southwest of Talca. The depth of the earthquake hypocenter was 

22 mi (35 km).

© Google, Europa Technologies, DMapas, and Inav/Geosistemas SRL

Figure 3. Map. Location of epicenter, seismic recording stations, and TIRT sites.
8

Chile has a long history of severe earthquakes along its approximately 1,860-mi (3,000-km)

coastline. Since the beginning of the 20th century, five earthquakes of M8.0 or higher have 

been recorded in Chile: M8.2 in 1906, M8.2 in 1943, M8.0 in 1985, and an M8.2 foreshock that 

preceded an M9.5 earthquake in 1960. Since 1973, there have been 13 events of M7.0 or greater.

The February 27, 2010, earthquake originated about 140 mi (230 km) north of the source region 

(the point of initial rupture) of the M9.5 earthquake of 1960, the largest instrumentally recorded 

earthquake in the world. The 2010 earthquake was approximately 190 mi (300 km) south of the 

source region of the M8.2 earthquake of 1906 and approximately 540 mi (870 km) south of the 

source region of the M8.5 earthquake of 1922.

Figure 4 shows the spatial relationship between previous earthquakes (M6.5 or greater, 1900–1963

and M5.5 or greater, 1964–present) and the 2010 M8.8 offshore Maule earthquake (depicted by a 

yellow star) and its associated aftershocks (depicted by yellow circles). The points of initial rupture

for the M8 or larger earthquakes are denoted by red circles with thick black borders. The estimated

rupture zones of the 1922 M8.5 and 1960 M9.5 earthquakes are marked in white. These rupture zones 

were defined by the areal extent of aftershocks or by geological and macroseismic observations. 

The 2010 M8.8 earthquake ruptured the portion of the South American subduction zone between 

these massive historical earthquakes, with the 1906 and 1922 megaquakes to the north and the 1960

megaquake to the south. The offshore Maule earthquake rupture zone was over 370 mi (600 km)

long and 81 mi (130 km) wide. The last time this portion of the subduction zone ruptured may 

have been the 1835 event experienced by Charles Darwin.

Figure 4. Map. Historic seismicity for Chile.

(1)
9

The 2010 offshore Maule earthquake occurred as thrust faulting on the interface between two plates,

with the Nazca plate moving down and landward below the South American plate. The slip rate 

between the two plates is estimated to be approximately 2.7 inches/year (70 mm/year).

(3) This rate

is much higher than the slip rate for the Cascadia subduction zone located along coastal Washington

and Oregon (approximately 1.4 inches/year (35 mm/year)), which explains the more frequent 

occurrence of this type of event along the Chilean coastline.(4) Ruegg et al. estimated that at least 

33 ft (10 m) of slip deficit had accrued on this segment of the plate boundary since the last time 

this portion of the subduction zone ruptured in 1835.

(3) GPS measurements indicate that the actual

horizontal movement of the overriding plate was as much as 12 ft (3.7 m) to the southwest at 

Concepción, 15 ft (4.7 m) at Constitución, and approximately 2 ft (0.5 m) at Santiago.

This type of earthquake is illustrated conceptually in figure 5. Overall, a tectonic plate descends, or 

“subducts,” beneath an adjoining plate in a stick-slip fashion. Between earthquakes, the plates slide 

freely at great depth, but at shallow depth, they stick together. A key aspect of this type of earthquake

is the buildup of stress at the stuck portion of the boundary. As the upper plate attempts to override

the lower plate, it deforms and is uplifted until the stress is too great. At that point, the stuck

portion of the boundary catastrophically ruptures, allowing the front portion of the upper plate to 

spring seaward and upward while the rest of the upper plate relaxes, causing subsidence. Field 

observations reported by GEER indicate that the coastline from the vicinity of the epicenter south

to the Arauco peninsula experienced uplift up to approximately 7 ft (2 m), whereas areas north of 

Constitución to Bucalemu experienced subsidence of up to about 3 ft (1 m).(2) Observations made

by TIRT at the sites investigated appear to confirm these findings. This type of coseismic

deformation has historically been accompanied by the generation of tsunamis. For the offshore 

Maule earthquake, a tsunami that severely impacted the coast north of the epicenter was produced.

As the area to the north of the epicenter also experienced significant coseismic subsidence, tsunami 

run-up was increased, resulting in greater onshore damage.

Figure 5. Illustration. Conceptual diagram of a subduction zone earthquake.

(2,4)

Another key aspect of this type of earthquake is the tendency of such an event to produce many 

strong aftershocks as the plates adjust to the sudden change in stresses following rupture. The 

aftershocks themselves can produce a significant hazard to the structures and facilities damaged 

by the main shock, as well as to the people within or near the partially damaged structures.

Figure 4 and figure 6 show maps illustrating the number and distribution of aftershocks, some of 

which approached M7.0. The cross section of the aftershock locations as a function of depth and 

distance in figure 6 illustrates the location of the rupture surface. The hypocenter of the main 

shock is marked with a star.
10

1 km = 0.621 mi

Figure 6. Illustration. Offshore Maule earthquake aftershock distribution with depth.

(5)

2.3 GROUND MOTIONS

The actual ground motions experienced at a particular site depend on the soil type and density 

present, as well as the three-dimensional aspects of the local geologic structure beneath and around

the site. In general, as soils become softer or looser, the ground motions tend to increase, especially

at longer periods. Basin effects can also amplify ground motions at a particular site. As the team 

traveled throughout the region impacted by the earthquake, it was apparent that damage to

transportation and other structures was concentrated in specific areas, with areas in between showing

little or no damage. GEER reported similar observations and attributed the difference to the 

presence of softer silts and clays at the affected sites and, in some cases, to basin effects.(2)
11

At the time of this report, the available ground motion data from the offshore Maule earthquake was

limited, and ground motion records were only available from some of the stations identified in

figure 3, shown as small rectangles with blue accelerograms. White icons denote stations where no 

accelerogram was available. Yellow icons denote stations where an accelerogram was available but 

digital records of the ground motion were not available. Pink icons (only one, in Concepción) 

denote stations where digital records of the ground motion were available.

The available records, summarized in table 2, are uncorrected and may be influenced by the 

structures in which the measurements were taken. The ground motion records are from two 

different sources, Boroschek, et al. and the University of Chile Department of Geophysics.(6,7)

Boroschek, et al. indicated that all ground motion monitoring equipment was in one-story structures,

but in some cases, other types of structures were nearby.(6) All of these records require careful

evaluation due to potential soil-structure interaction effects.(6,7) Specifically, Boroschek, et al. noted

that the high accelerations at the Maipú site were being reviewed for soil-structure interaction 

effects. Soil classifications at each of these monitoring sites were being investigated at the time 

of this report.(6) Boroschek, et al. also cautioned that the quake data recorder (QDR) only recorded 

the first 100 s of motion.(6) However, the motion actually lasted about 140 s. Therefore, the last 

40 s of motion is missing from these records. QDR equipment was used for all of the ground

motion records in table 2, except Universidad de Chile Deptartamento de Ingeniería Civil (interior 

building); Estación Metro Mirador, Santiago; and Colegio San Pedro, Concepción. Examples of 

these ground motions and their respective response spectra (where available) are provided in 

figure 7 through figure 14.

Table 2. Summary of available earthquake records for the offshore Maule earthquake.(6)

Location of Monitoring Station

Source of 

Information

Maximum 

Horizontal 

Acceleration 

(g)

Maximum 

Vertical 

Acceleration 

(g)

Universidad de Chile, Depto.

Ing. Civil (interior building) RENADIC(6) 0.17 0.14

Estacion Metro Mirador, Santiago RENADIC(6) 0.24 0.13

CRS Maipú RM RENADIC(6) 0.56 0.24

Hospital de Tisné RM RENADIC(6) 0.30 0.28

Hospital Sótero de Río RM RENADIC(6) 0.27 0.13

Hospital de Curicó RENADIC(6) 0.47 0.20

Hospital de Valdivia RENADIC(6) 0.14 0.05

Vina del Mar, Marga Marga RENADIC(6) 0.35 0.26

Vina del Mar (Centro) RENADIC(6) 0.33 0.19

Colegio San Pedro, Concepción

Red Sismológica 

Nacional

0.65 NS, 

0.58 EW 0.60

San Pedro de la Paz, Colegio 

Concepción USGS(8)

0.65 NS, 

0.61 EW 0.58

NS = North-south.

EW = East-west.
12

Figure 7. Illustration. Accelerogram from Universidad de Chile, Depto. Ing. Civil

(interior building).

(6)

Figure 8. Illustration. Response spectra from Universidad de Chile,

Depto. Ing. Civil (interior building).

(6)
13

Figure 9. Illustration. Accelerogram from CRS Maipú RM.

(6)

Figure 10. Illustration. Response spectra from CRS Maipú RM.

(6)
14

Figure 11. Illustration. Accelerogram from Hospital de Curicó.(6)

Figure 12. Illustration. Response spectra from Hospital de Curicó.(6)
15

Figure 13. Illustration. Accelerogram from San Pedro De La Paz, Colegio Concepción.

(8)

Figure 14. Illustration. Response spectra from San Pedro De La Paz, Colegio Concepción.

(8)

East-West Direction

Up-Down Direction

PGA = 0.61g Horiz.

PGA = 0.58g Vert.

North-South Direction

PGA = 0.65g Horiz.

Sa for 5% damping

East-West Direction

Sa for 5% damping

Up-Down Direction
16

Figure 13 is the only time history for which a digital file was available at the time of this report.

For the east-west direction in that time history, the bracketed duration (i.e., A > 0.05 g) was 152 s

and the duration of significant shaking based on normalized Arias intensity was approximately 

76 s. Visually and in terms of peak values, this ground motion record is very similar to the one 

obtained at Colegio San Pedro. Figure 15 through figure 19 are maps marked with accelerograms

and TIRT structure sites to show the location of the ground motion recording stations relative to 

the structure sites visited. The symbols used in figure 3 are also used in these figures. See chapter 

1 and chapter 3 for a more detailed description of each structure site.

© Google, Inav/Geosistemas SRL, DMapas, and GeoEye

Figure 15. Map. Locations of ground motion sensors and structures visited by TIRT—

Santiago vicinity.
17

© Google, Inav/Geosistemas SRL, DMapas, GeoEye, and DigitalGlobe

Figure 16. Map. Locations of ground motion sensors and structure visited by TIRT—

Santiago to Rancagua.
18

© Google, Cnes/Spot Image, DMapas, GeoEye, and DigitalGlobe

Figure 17. Map. Locations of ground motion sensors and structures visited by TIRT—

Curicó, Talca, Iloca vicinity.
19

© Google, Inav/Geosistemas SRL, DMapas, and GeoEye

Figure 18. Map. Locations of ground motion sensors and structures visited by TIRT—

Concepción vicinity.
20

© Google, DigitalGlobe, DMapas, and GeoEye

Figure 19. Map. Locations of ground motion sensors and structures visited by TIRT—

Maule epicenter to Tubul vicinity.

GEER provided geologic information that can be used in combination with the subsurface 

information gathered by TIRT to preliminarily assess the potential effect subsurface conditions 

may have had on ground motion severity.(2) The subsurface information gathered by TIRT is

reported in appendix A. For example, the GEER report indicates that relatively soft silts and clays

and, in some cases, ash deposits are likely to be encountered in the western sector of Santiago, 

where structure sites 1–6 are located.(2) A boring log obtained for the original bridge design at 

sites 1 and 2 indicates that subsurface soils there consist of 16 ft (5 m) of soft to medium stiff 

clay underlain by more than 33 ft (10 m) of stiff to very stiff silty clay with layers of fine sand.

Based on the AASHTO and National Earthquake Hazards Reduction Program (NEHRP) seismic 

site classification systems, these soils would likely be classified as Site Class E soils, and significant

ground motion amplification would be expected. The eastern sector of Santiago, where most of the

available earthquake ground motion measurements were taken, generally consists of sands and 

gravels. It could therefore be expected that, in general, ground motion measurements in the eastern 

sector of Santiago should be less severe than those obtained in the western sector because of better

soil conditions in addition to being farther away from the earthquake epicenter. A comparison of

the ground motions in figure 7 with those in figure 9 appears to support this conclusion—the 

peak accelerations in figure 7 (0.17 g) are considerably less than those in figure 9 (0.56 g). The 

other ground motion records obtained in the eastern sector of Santiago range from 0.24 g to 0.30 g,

still considerably lower than the peak measured at CRS Maipú RM (see figure 9). The response 
21

spectra also seem to indicate the potential effect of softer soils in that the peak spectral acceleration

is pushed out to longer periods and the spectral acceleration magnitude is two to four times higher

than that of the eastern sites. Based on the location of the first six sites, a reasonable preliminary

conclusion is that the acceleration time history and response spectra for CRS Maipú RM, shown 

in figure 9 and figure 10, are most representative of the ground motions that occurred at the first 

six bridge sites. Once subsurface geotechnical data for these earthquake ground motion monitoring

sites becomes available and the soil-structure interaction effect is fully investigated, these 

preliminary conclusions can be confirmed.

The next set of structure sites, sites 7–10, are probably on the margin of the poorer soil conditions

based on the geologic data published by GEER.(2) However, some geotechnical boring information

is available at sites 7, 8, and 10 indicating the presence of at least 33 to 49 ft (10 to 15 m) of 

medium to stiff silty clays or sandy silts, with stiff to hard silts and clays and some dense sand or 

gravel layers at depth (see appendix A). Based on the AASHTO and NEHRP classification systems,

these soils would likely be classified as Site Class D or E soils, and significant ground motion

amplification would be expected. Again, the CRS Maipú RM ground motions (shown in figure 9

and figure 10) appear to be the most applicable to these structure sites with due consideration 

given to the cautionary issues previously discussed.

Sites 11–13 and 32 are far enough away from Santiago that closer ground motion records should 

be used. The only ground motion record available in this area is the one at Curicó (see figure 11

and figure 12). The GEER report does not provide specific geologic information, but deposits at 

Talca (south of Curicó) are likely to be alluvial and marine sediments and are probably highly 

variable.(2) Therefore, the strength of the ground motions in this region is likely to be highly 

variable, as indicated by the inconsistent damage to structures and facilities.

South of Talca and closer to the coast, the ground motion records in the interior may not be

representative of the ground motions experienced, considering proximity to the epicenter and

subsurface conditions present. Fortunately, two ground motion records are available in Concepción

(see figure 13). According to GEER, Concepción has both poor soils (loose alluvial and marine 

deposits) and a geologic structure, due to major faults running through the area, that could contribute

to amplification of ground motion (see figure 20 and figure 21).

(2) Damage due to the offshore

Maule earthquake was particularly severe in Concepción. The spatial distribution of the damage 

is apparently similar to the damage seen during the 1960 M9.5 event, trending along parallel linear

zones. This similarity shows the effect geologic basin structure and soil deposits may have had on 

the ground motions. The location of the Colegio San Pedro monitoring station is shown in 

figure 21. This monitoring station is located in between two significantly damaged areas. Therefore,

it is not clear if the Colegio San Pedro ground motion record, which is similar to the San Pedro de

la Paz record provided in figure 13, represents the most damaged areas. Considering the high peak

acceleration of 0.65 g in the horizontal direction and 0.60 g in the vertical direction, the long 

duration, and the similarity to the San Pedro de la Paz ground motion, a reasonable preliminary 

conclusion is that these two ground motion records are representative of what occurred in the 

significantly damaged areas. While it is unknown if there are any geologic basin issues for the 

sites farther south (e.g., sites near Tubul), the soil conditions are obviously poor and probably 

similar to conditions in parts of Concepción (see chapter 5). Therefore, this ground motion 

should be considered applicable to sites 17–29.
22

1 km = 0.621 mi

Figure 20. Illustration. Geologic cross section of Concepción.

(2)

Figure 21. Map. Damaged sections in downtown Concepción.

(2)

Not much information is available on the subsurface conditions of sites 15, 16, and 30–32, and there 

are no nearby ground motion records for those locations. Until more is known, a compromise 

between the Curicó and Concepción ground motions could be used for preliminary analyses.

Once more ground motion records and geotechnical subsurface data are available to confirm these

conclusions (based primarily on geologic mapping), the ground motions most applicable to each 

structure site can be better assessed.
23

CHAPTER 3. OVERVIEW OF BRIDGE PERFORMANCE AND SEISMIC DESIGN 

REQUIREMENTS IN CHILE

3.1 OVERVIEW OF BRIDGE PERFORMANCE

3.1.1 Damage Statistics

Of the nearly 12,000 bridges in Chile, about 200 were damaged in the offshore Maule earthquake,

including 20 structures with one or more collapsed spans. These 12,000 bridges include about

4,750 culverts and pedestrian overcrossings and 7,250 highway bridges. Of the highway bridges, 

6,800 are publicly owned by MOP and 450 are owned by private companies called concessions, 

which have designed and constructed several major toll roads in Chile. Table 3 shows some 

damage statistics.

Table 3. Number of damaged bridges.

Owner

Number of 

bridges1

Number of 

damaged 

bridges2

Number of 

collapsed 

bridges3

MOP 6,800 103 10

Concessions 450 100 8

Total 7,250 203 18 1 Excluding culverts and pedestrian overcrossings. 2 Including bridges with collapsed spans. 3 Bridges with one or more collapsed spans.

3.1.2 Bridges Visited By Reconnaissance Team

As noted in chapter 1, TIRT visited 41 damaged bridges at 32 sites in Chile from Santiago to 

Arauco over a 9-day period from April 4 to 13, 2010. Table 4 provides a summary of the bridges 

visited along with brief notes about the observed damage. The locations of these sites are shown 

in figure 1 and figure 3.
24

Table 4. Bridges visited by reconnaissance team.

Site

No. Site Name Location

Year 

Built

Structure Description

Type Geometry Damage

1a

Américo Vespucio/ 

Miraflores eastbound

Santiago

3 spans, PC-PSC, 5-column bents, seat abutment, MSE walls, tie downs Skewed Collapsed

1b

Américo Vespucio/ 

Miraflores westbound 3 spans, PC-PSC, 5-column bents, seat abutment, MSE walls, tie downs Skewed Collapsed

2a

Américo Vespucio/

Lo Echevers eastbound

Santiago 2004

3 spans, PC-PSC, 5-column bents, seat abutment, MSE walls, tie downs Skewed Collapsed

2b

Américo Vespucio/

Lo Echevers westbound 3 spans, PC-PSC, 5-column bents, seat abutment, MSE walls, tie downs Skewed Moderate

3 I-5/14 de la Fama Santiago 15 spans, PC-PSC, 5-column bents, seat abutment, MSE walls, tie downs Straight Moderate

4

Pedestrian bridge over 

Route 5 Santiago 3 spans, 2 SG, single columns Straight Collapsed

5

Quilicura railway crossing 

at Avenida Manuel Antonio 

Matta Santiago 3 spans, 5 SG, 5-column bents, seat abutment, CIP walls Skewed Collapsed

6a

Américo Vespucio/

Independencia westbound

Santiago 2004

6 spans, PC-PSC, 5-column bents, seat abutment, MSE walls, tie downs Straight Severe

6b

Américo Vespucio/

Independencia eastbound

4 spans, PC-PSC, flared bents, diaphragms, seat abutment, MSE walls, 

seismic bars Straight Moderate

6c

Américo Vespucio/

Independencia westbound 

exit ramp 4 spans, PC-PSC, 1-column bent, seat abutment, MSE walls, tie downs Curved Minor

6d

Américo Vespucio/

Independencia westbound 

entrance ramp 3 spans, RC box girder, single-column bents, seat abutment, CIP walls Curved Minor

7

Avenida Romero Acceso 

Sur overpass Paine 2001

2 spans, PC-PSC, 4-column bents, seismic bars, seat abutment, earth 

embankment Skewed Collapsed

8

Avenida Chada Acceso Sur 

overpass Paine 2001

2 spans, PC-PSC, 4-column bents, seismic bars, seat abutment, earth 

embankment Straight Severe

9 Maipú River Buin 1970 13 spans, RC girders, A-shaped RC piers, diaphragm Straight Severe

10a

Route 5 railway crossing at

Hospital westbound

Buin 2001

2 spans, PC-PSC, 3-column bents, seat abutment, tie downs Skewed Collapsed

10b

Route 5 railway crossing at

Hospital eastbound 2 spans, PC-PSC, wall-pier bents, seat abutment, diaphragms, tie downs Skewed Minor

11

Estribo Francisco Mostazal 

(Avenida Independencia) Santiago 2001 1 span, PC-PSC, seat abutment, seismic bars Skewed Minor

12

Las Mercedes Route 5 

overpass Rancagua 2001

2 spans, PC-PSC, 2-column bents, seismic bars, seat abutment, earth 

embankment Straight Severe

13a Claro River San Rafael 1870 7 spans, brick masonry arch (1870) Straight Collapsed

See notes at end of table.
25

Table 4. Bridges visited by reconnaissance team—Continued.

Site

No. Site Name Location

Year 

Built

Structure Description

Type Geometry Damage

13b Claro River San Rafael 5 spans, RC arch, approach spans seat abutment Straight Minor

14 Pichibudis Iloca 1 span, 2 SG, seat abutment Straight Moderate

15 Mataquito Iloca 2008 4 spans, PC-PSC, 3-column bents, seat abutment, diaphragm, seismic bars Straight

16

Cardenal Raúl Silva 

Henríquez Constitución 2002 22 spans, 3 SG, multicolumn bents, seat abutment Straight Moderate

17 Llacolen Concepción 2000 Long bridge, PC-PSC, multicolumn bents, seat abutment Straight Collapse

18

Chepe railroad bridge over 

Biobío River Concepción

1889, 

retrofitted 

in 2005 Long bridge, truss, steel-pipe column supports, CIP walls Straight Moderate

19a

Puerto de Coronel Muelle 

Norte

Coronel

Wharf with steel pile bents and RC deck Straight Moderate

19b

Puerto de Coronel Muelle 

Sur Wharf with steel pile bents, RC deck, LRB base isolation system Straight None

20 Raqui 1 Raqui 2 spans, SG, wall-type bents, seat-type abutment Straight Minor

21 Raqui 2 Raqui 4 spans, SG, wall-type intermediate bent, seat-type abutment Straight Collapsed

22 Tubul Raqui 8 spans, 3 SG, wall pier bents, seat abutments, diaphragm, seismic bars Straight Collapsed

23 El Bar Arauco 1 span, 2 SG, seat abutments, retrofitted bridge Straight Moderate

24a Ramadillas (west, old)

Arauco

14 spans, pier-wall bent, 4 steel beams, seat abutment Straight Collapse

24b Ramadillas (east, new) 8 spans, 3 SG, wall pier bents, seat abutment, diaphragm, seismic bars Straight Settlement

25 Juan Pablo II Concepción 1964 Multispan bridge, PC-PSC, multicolumn bents, seat abutment Straight Settlement

26 Biobío River (old) Concepción 1943 Multispan bridge, 3 steel beams, pier-wall bents, seat abutment (1943) Straight Collapsed

27 La Mochita Concepción 2005 4 spans, PC-PSC, 2-column bents, seat abutment, seismic bars Straight Severe

28

Vía Elevada 21 de 

Mayo/Cruce Ferroviario Concepción 2 spans, RC beams, multicolumn bent, seat abutments Skewed Collapsed

29 Rotonda General Bonilla Concepción 2010 5 spans, 5 SG, 6-column bents, seat abutments Straight Minor

30 Itata River Coelemu 1990 22 spans, 2 SG, pier-wall bents, seat abutments Straight Moderate

31 San Nicolás San Nicolás 2 spans, PC-PSC, pier-wall bent, seat abutments Straight Severe

32 Muros Talca (SW) Talca 1 span, PC-PSC, seat abutment, diaphragm, seismic bars Skewed Moderate

LRB = Lead-rubber bearing

PC-PSC = Precast prestressed concrete I-girders.

SG = Steel girders.

RC = Reinforced concrete.

MSE = Mechanically stabilized earth walls.

CIP = Cast-in-place RC walls.

Note: Blank cells indicate unavailable data.
26

3.1.3 Damage Summary

The following principal types of damage were observed:

Span unseating due to lack of transverse diaphragms and shear keys in precast pretressed 

concrete superstructures.

Span unseating due to skew and insufficient support length.

Span unseating due to ground movement caused by liquefaction-induced spreading.

Girder distortion, column damage, and heavy scour due to tsunami wave forces.

Column failure due to ground movement caused by liquefaction-induced spreading.

Detailed descriptions of the damage types are given in chapter 4 and chapter 5. 

Many of the bridges built by concessions used precast prestressed concrete girder superstructures

without diaphragms or shear keys for transverse restraint. Vertical rods called seismic bars and 

hold-down ties were used to prevent uplift after high vertical ground accelerations were recorded 

during the 1985 earthquake. These rods and ties were largely ineffective in the transverse direction,

and many spans slid sideways on their cap beams. This lack of restraint also allowed a number of 

two-span bridges to rotate about a vertical axis through the pier and slide off their abutment seats.

In addition, several skewed spans with diaphragms and shear keys rotated about a vertical axis 

and were unseated in their acute corners due to insufficient support length. Straight bridges built 

before the concession era and those with cast-in-place (CIP) diaphragms and concrete shear keys 

performed well.

Despite higher than anticipated spectral accelerations, column damage was slight, perhaps because

the lack of transverse restraint and insufficient support length allowed many superstructures to 

separate from their substructures, limiting the demand on the columns. When the superstructure did 

not separate, column damage was more likely to occur, such as with the shear failures due to 

imposed displacements from liquefaction-induced lateral spreading in several columns under the 

approach spans to the Juan Pablo II bridge across the Biobío River in Concepción.

In addition to this bridge, liquefaction-induced lateral spreading or settlement is believed to be 

responsible for the collapse or serious damage of many other structures along the coast, including

the Llacolen, Chepe, Ramadillas, and Tubul bridges.

Bridges on coastal highways also sustained tsunami damage, such as the lateral distortion of the 

superstructure of the Pichibudis bridge just north of Iloca, the undermining of several piers due 

to scour, and the puncture of steel pile bents by floating debris in the Cardenal Raúl Silva 

Henríquez bridge across the Maule River at Constitución. 
27

3.2. SEISMIC DESIGN REQUIREMENTS FOR BRIDGES

In Chile, as in other highly seismic countries, earthquake engineering has evolved over time, and 

advancements can be linked to the occurrence of large earthquakes. 

According to Rodrigo Flores, the first step toward modern seismic design in Chile occurred after the 

1906 Valparaiso earthquake, when the government created the Chilean Seismological Institute.

(9)

After the 1928 Talca earthquake, another important step was taken with the passage of the 1931 

Act of Construction and Urbanization, which established basic requirements for the seismic design

of buildings. This document evolved over time until the 1972 creation of the Chilean Seismic Code

for Buildings, which was based on U.S. and Japanese seismic codes. After the 1985 earthquake, 

studies were conducted to develop uniform risk maps for the country and three seismic zones were

established, with the highest risk occurring along the Pacific Coast. This zoning was reflected in 

a 1996 update of the seismic building code.

Seismic design methodologies for bridges have also been based on U.S. and Japanese experience.

According to Alex Unión and Rodolfo Saragoni, the design of concrete bridges in Chile before 

1950 was based on the handbook published by Alberto Claro Velasco, Normas para el Cálculo y 

Proyecto de Puentes Carreteros de Hormigón Armado (Standards for the Design and Protection 

of Reinforced Concrete Road Bridges). After the mid-1950s, most designs were based on the 

AASHTO Standard Specifications for Highway Bridge Design.

(10)

Before the mid-1980s, the seismic design coefficient for bridges was 0.12. This coefficient was 

increased to 0.15 following the 1985 earthquake, and a modified version of Division I-A of the 

AASHTO Standard Specifications was adopted in 1998. The design coefficient was not changed 

until 2001 when three seismic zones were introduced with peak ground accelerations (PGAs) of 

0.2, 0.3, and 0.4 g (see figure 22). In addition, the soil factors were modified along with the response

modification factors, and an allowance for the effect of scour was introduced. Column design was

required to be in accordance with the AASHTO requirements for Seismic Performance Categories C

and D of Division I-A.(10) These provisions can be found in section 3.1004 of the MOP Manual 

de Carreteras (Highway Handbook) and are summarized in appendix B.

(11)
28

Figure 22. Map. Seismic zone map for central Chile.

(11)
29

CHAPTER 4. STRUCTURAL PERFORMANCE OF HIGHWAY BRIDGES

4.1 OVERVIEW

General observations of damage to highway bridges are reported in this chapter. The bridges are 

organized according to bridge type, such as concrete and steel girder superstructures, arches, and 

trusses. They are further divided into regular and irregular structures in terms of skew angle and 

curvature. The bridge site numbers are provided for reference.

The following bridges are discussed in section 4.2, Performance of Straight Concrete Bridges:

Independencia (sites 6a–6d).

Avenida Chada (site 8).

Las Mercedes (site 12).

Llacolen (site 17).

Juan Pablo II (site 25).

Ramadillas (sites 24a and 24b).

The following bridges are discussed in section 4.3, Performance of Skewed and Curved Concrete 

and Steel Bridges: 

Miraflores (sites 1a and 1b).

Lo Echevers (sites 2a and 2b).

Avenida Romero (site 7).

Route 5 railway overcrossing at Hospital (sites 10a and 10b).

Matta Quilicura (site 5).

The following bridges are discussed in section 4.4, Performance of Straight Steel Bridges: 

Tubul (site 22).

Cardenal Raúl Silva Henríquez (site 16).

Biobío River (old) (site 26).

Pichibudis (site 14).

El Bar (site 23).
30

Itata (site 30).

Pedestrian bridge over Route 5 (site 4).

Lastly, the following bridges are discussed in section 4.5, Performance of Other Bridge Types:

Claro River (sites 13a and 13b).

Chepe railroad bridge over Biobío River (site 18).

Maipú River (sites 9a–9c).

4.2 PERFORMANCE OF STRAIGHT CONCRETE BRIDGES

4.2.1 Américo Vespucio/Independencia

Two bridges oriented in the east-west direction and two ramps connected to westbound traffic 

were inspected at this site. The eastbound bridge, built in 2004, is a four-span structure with five 

discontinuous precast prestressed girders and a continuous deck. It almost collapsed during the 

earthquake, experiencing a lateral offset of greater than 1.6 ft (0.5 m), as indicated by the use of 

temporary heavy supports shown in figure 23 through figure 25. The superstructure had steel 

stoppers but no intermediate or end diaphragms. As shown in figure 24 and figure 25, both the 

curtain wall of the abutment and the steel stoppers were significantly damaged. Each stopper was 

anchored with two bolts embedded in the cap beam at a depth of 35 inches (900 mm), as shown 

in figure 26. Steel stoppers were mainly used to restrain the vertical movement of girders.

Figure 23. Photo. Damage to eastbound Independencia bridge.

Figure 24. Photo. Curtain wall at east abutment of eastbound Independencia bridge.
31

Figure 25. Photo. Damage to steel stoppers on eastbound Independencia bridge.

Source: MOP

Figure 26. Illustration. Details of steel stoppers.

The westbound bridge, built in 1997, is a four-span structure with six discontinuous precast 

prestressed girders and a continuous deck. Each bent has two flared wall piers, as shown in 

figure 27. The bridge was designed with both intermediate and end diaphragms. The bridge 

survived the earthquake with excessive deformations in the seismic bars and damage to concrete

shear keys at abutments and intermediate bents (see figure 28 through figure 30).
32

Figure 27. Photo. Typical flared wall pier of westbound Independencia bridge.

Figure 28. Photo. Displaced seismic bar on westbound Independencia bridge.

Figure 29. Photo. Shear key damage at abutment of westbound Independencia bridge.

Figure 30. Photo. Shear key damage at intermediate bents of westbound

Independencia bridge.

In comparison to the eastbound bridge, the presence of diaphragms seems to have helped maintain

the structural integrity of the westbound bridge. More importantly, the concrete shear keys of the 

westbound bridge served their design purpose during the earthquake, as opposed to the steel stoppers

in the eastbound bridge. No visible foundation damage was observed for either structure.

Additional views of the shear key damage in the westbound bridge and its entrance ramp are shown

in figure 31 through figure 33. As shown in figure 31, shear key failures occurred in almost the 

entire structure. The entrance ramp next to the bridge is supported on two hammerhead piers at 

Seismic bars
33

the intermediate bents, as shown on figure 32. The shear keys at the abutment were damaged. 

Figure 33 shows an approximately 1.5-inch (38-mm)-wide gap between the abutment and the 

surrounding soil caused by strong ground shaking.

Figure 31. Photo. Westbound Independencia bridge.

Figure 32. Photo. Entrance ramp to westbound Independencia bridge.

Figure 33. Photo. Shear key damage at entrance ramp abutment.

The exit ramp for westbound traffic remained in service after the earthquake. Minor damage due 

to pounding in the corner of the beam recess was observed on a flared wall pier close to the 

westbound bridge, as shown in figure 34 and figure 35.

Figure 34. Photo. Flared wall pier of exit ramp from westbound Independencia bridge.
34

Figure 35. Photo. Close-up of damage to exit ramp from westbound Independencia bridge.

4.2.2 Avenida Chada Acceso Sur Overpass

The two-span Chada overpass, shown in figure 36 through figure 39, is oriented approximately 

east-west. The bridge superstructure was displaced significantly and rotated counterclockwise in 

plan during the earthquake. The transverse offset between the deck and abutment seats ranged from

25 to 30 inches (640 to 780 mm). The longitudinal gap between the end of girders and the abutment

was 3.5 inches (90 mm). The bottom flange and web of the exterior precast girder that hit the 

curtain walls at each abutment failed in shear, as shown in figure 36. The failure of the bottom

flange can be seen in figure 37. The two curtain walls at the intermediate bent were both damaged,

as shown in figure 38. The bridge experienced strong shaking, indicated by an induced gap around 

the columns, as shown in figure 39. No structural damage was observed in the bridge columns.

Figure 36. Photo. Curtain wall shear failure at abutment of Chada bridge.

Figure 37. Photo. Bottom flange damage to Chada bridge.
35

Figure 38. Photo. Curtain wall damage at bent of Chada bridge.

Figure 39. Photo. Soil separation from column of Chada bridge.

Although the bridge is not skewed, significant rotation was observed in its superstructure, as 

illustrated in figure 40. In the absence of restraint at the abutments (i.e., no diaphragm), the period 

of the rotational mode of vibration was shorter than the translational modes and appeared to have

been strongly excited in the earthquake either by the translational components of the ground 

motion or by the rotational component. If the former, some degree of eccentricity between the 

centers of mass and stiffness would need to be present, which is quite likely given the slumping 

of the abutment fills and the duration of shaking.

1 cm = 0.39 inches

Figure 40. Illustration. Plan view of superstructure rotation of Chada bridge.

4.2.3 Las Mercedes Route 5 Overpass

The two-span Las Mercedes overpass is oriented approximately east-west and is a concrete-girder

structure with a slight skew, as shown in figure 41. The superstructure comprises three precast 

prestressed girders and is supported on a two-column intermediate bent and two seat-type 

abutments, all resting on drilled shafts. Two seismic bars (deformed rebar) were installed 

between pairs of adjacent girders at all supports. Like the Chada bridge, the Las Mercedes bridge 

78

64

cm

cm
36

experienced a significant counterclockwise rotation during the earthquake. The exterior girders 

rotated off their abutment seats at both ends, resulting in a longitudinal crack between the 

exterior and interior girders in the deck slab.

Figure 41. Photo. Las Mercedes bridge and girder unseating at abutments.

In addition, one transverse crack across the bridge deck was observed at the intermediate bent. 

The curtain walls were severely damaged when struck by the bridge superstructure. Temporary 

supports were provided for the unseated exterior girders, as shown in figure 41.

4.2.4 Llacolen Bridge Over Biobío River

The Llacolen bridge, a major crossing over the Biobío River in Concepción is oriented 

approximately northeast-southwest and is a multispan, simply supported concrete girder bridge.

Each span consists of a deck slab and six precast prestressed girders that are supported on two

five-column bents with an inverted-T cap beam. Two seismic bars are provided between each 

pair of adjacent girders. Figure 42 shows a general view of the bridge. 

Figure 42. Photo. Llacolen bridge looking south-southwest.

As shown in figure 43, a simply supported span in the eastern approach to the bridge dropped from

its seat at the river end of the span, and a temporary Bailey bridge was erected to give traffic access

to the main crossing. As shown in figure 44, the end bent supporting the unseated span remained 

intact except for concrete spalling underneath the cap beam, as shown in figure 45. The seismic 

bars shown in figure 44 were deformed almost 180 degrees as the span collapsed and they were 

East Biobío River

abutment West 

abutment
37

pulled from the diaphragms. The columns supporting the cap beam experienced flexural cracks 

at the level of the rock rip rap on the bank below the bridge (see figure 46 and figure 47). The

cracks appeared on the river side of the columns, where tension developed as the superstructure 

held the columns against the lateral movement imposed on the foundations by liquefaction-induced

lateral spreading of the banks. The bent and columns at the opposite end of the unseated span 

were undamaged. However, as shown in figure 48, the nearby ground settled up to 1.3 ft (0.4 m)

and experienced significant shaking, resulting in a 0.82-ft (0.25-m) separation between the 

columns and their surrounding ground.

Figure 43. Photo. Unseated simply supported span in eastern approach to Llacolen bridge.

Figure 44. Photo. Abutment at unseated end of span in eastern approach to Llacolen bridge.

Figure 45. Photo. Concrete spalling beneath cap beam of eastern approach to Llacolen bridge.

Seismic bars

Concrete spalling

Deformed 

Seismic Bars

Bailey bridge
38

Figure 46. Photo. Flexural cracks at the level of rip rap 

on eastern approach to Llacolen bridge.

Figure 47. Photo. Close-up view of cracks at the level of rip rap 

on eastern approach to Llacolen bridge.

Figure 48. Photo. Ground settlement and lateral movement 

of eastern approach to Llacolen bridge.

The entrance ramp for westbound traffic is a three-span simply supported box girder bridge. The 

girders are supported on wall piers with neoprene pads and fitted with seismic bars. The ramp 

was almost unseated at the river end of the last span, as shown in figure 49 and figure 50. At one 

intermediate bent, the seismic bars were rusty and the neoprene pads had degraded, as shown in 

figure 51 through figure 53.

Settled 1.3 ft (0.4 m); 

separated 0.82 ft (0.25 m)
39

Figure 49. Photo. Exterior face of westbound ramp to Llacolen bridge at the river end.

Figure 50. Photo. Interior face of westbound ramp to Llacolen bridge at the river end.

Figure 51. Photo. Bent of westbound ramp to Llacolen bridge.

Figure 52. Photo. Seismic bar condition on westbound ramp to Llacolen bridge.

Seismic bar
40

Figure 53. Photo. Neoprene pad degradation on westbound ramp to Llacolen bridge.

At the west end of the Llacolen bridge, soil settlement and liquefaction were observed at multiple 

locations near the bridge bents, as shown in figure 54 and figure 55. Pounding-induced damage was

also observed at the supports of precast girders, as shown in figure 56 and figure 57. Minor damage

also occurred when the girders engaged the concrete keys on the west abutment (see figure 58).

Figure 54. Photo. Columns with soil marks indicating ground settlement 

at west end of Llacolen bridge.

Figure 55. Photo. Sand boils near columns at west end of Llacolen bridge.

Figure 56. Photo. Spalling in deck slab at west end of Llacolen bridge.

Neoprene pad
41

Figure 57. Photo. Horizontal crack in web of end girders at west end of Llacolen bridge.

Figure 58. Photo. Spalling of shear key at southwest abutment of Llacolen bridge.

4.2.5 Juan Pablo II Bridge Over Biobío River

The Juan Pablo II bridge is another major crossing over the Biobío River in Concepción. It is 

oriented northeast-southwest. The northern approach to this bridge is part of an interchange that 

provides access to the bridge from a riverfront highway, as shown in figure 59. In addition to exit 

and entrance ramps, the interchange comprises a two-span approach structure and a two-span 

highway overpass. Each bridge superstructure consists of discontinuous precast prestressed

girders and a continuous deck.

Figure 59. Photo. Approaches to Juan Pablo II bridge.

Biobío River

Bridge over river- Two-span bridge front highway

Juan Pablo II Bridge
42

Although the riverfront overpass was generally undamaged, the two-span approach structure was 

damaged both globally and locally at the intermediate bent. As shown in figure 60 through figure 62,

the intermediate bent had settled with respect to the two end bents, leading to a substantial depression 

in the roadway surface. This intermediate bent comprises two rectangular columns that most likely

experienced uneven liquefaction-induced settlement during the earthquake (see figure 63). The 

consequential change in stiffness between the two columns shifted the longitudinal seismic force 

from one column to the other, which then failed in shear, as shown in figure 64. The transverse 

reinforcement of the failed column was spaced at 17.8 inches (457 mm). Figure 65 and figure 66

show the soil surface prior to the earthquake on the face of undamaged column and cracking in 

the cap beam above. 

Figure 60. Photo. Southwest view of northern approach to Juan Pablo II bridge, 

facing Biobío River.

Figure 61. Photo. Northeast view of northern approach to Juan Pablo II bridge, 

facing away from Biobío River.

Figure 62. Photo. Intermediate bent of northern approach to Juan Pablo II bridge.
43

Figure 63. Photo. Column settlement under approach to Juan Pablo II bridge 

(cracks on far side).

Figure 64. Photo. Reduced column height due to shear failure under approach 

to Juan Pablo II bridge.

Figure 65. Photo. Soil surface on far side of undamaged column under approach

to Juan Pablo II bridge.

Column settled with 

surrounding soils

Soil surface 

prior to earthquake

Column failed 

in shear
44

Figure 66. Photo. Cracks on far side of undamaged column under approach 

to Juan Pablo II bridge.

Consistent with the behavior of the approach structure, the spans of the main bridge also appeared

to have experienced uneven support settlement, which tilted the columns and rotated the bridge 

deck about the centerline of the bridge, as seen in figure 67. 

Figure 67. Photo. Differential settlement underneath first span over water 

at northern end of Juan Pablo II bridge.

In addition, lateral spreading of the northeast bank pushed the columns of the bent toward the river.

However, this movement was restrained by the superstructure, and the columns were consequently

placed in single shear. One failed catastrophically (see figure 68 through figure 70), and the other 

was seriously damaged (see figure 71). The difference in behavior between the two columns might

have been due to the uneven settlement and rotation of the bridge deck. The longitudinal movement

of the first column was approximately 21.8 inches (559 mm) at the top of the shear plane and 

15.8 inches (406 mm) at the bottom. Similar behavior was also observed in the Llacolen bridge, but 

the damage to the columns was not as severe.

Cracks

Bridge 

Settlements
45

Figure 68. Photo. Shear failure in upstream column at northern end of Juan Pablo II bridge.

Figure 69. Photo. Front face of failure plane at northern end of Juan Pablo II bridge.

Figure 70. Photo. Back face of failure plane at northern end of Juan Pablo II bridge.
46

Figure 71. Photo. Shear failure in downstream column at northern end 

of Juan Pablo II bridge.

4.2.6 Ramadillas Bridges

Two bridges cross the Piteateo River close to the town of Ramadillas in Arauco (see figure 72). 

The bridges are generally oriented north-south. The east bridge is a relatively new structure with 

total length of approximately 853 ft (260 m) comprising eight spans spaced at 105 ft (32 m). The 

superstructure is composed of three parallel precast prestressed girders simply supported between

piers and a reinforced concrete deck. The substructure is composed of two seat-type abutments and

seven intermediate wall-type piers (see figure 73 and figure 74). Reinforced concrete diaphragms 

are used at each abutment and intermediate pier with vertical restrainers and shear keys. This 

structure has two lanes, one for each direction of traffic.

© Google and GeoEye

Figure 72. Photo. Satellite image of Ramadillas bridges.
47

n

Figure 73. Photo. Side view of east Ramadillas bridge.

Figure 74. Photo. Wall-type pier beneath east Ramadillas bridge.

During the earthquake, the piers near the river settled due to liquefaction, causing vertical 

misalignment of the deck, as shown in figure 75. Sand boils appeared at the bottom of some 

piers, as shown in figure 76.

Figure 75. Photo. Deck misalignment under east Ramadillas bridge.

Settlement

RC Diaphragm

Wall-type pier Shear key

S. Abutment N. Abutment

Wall-type pier
48

Figure 76. Photo. Sand boils close to pier under east Ramadillas bridge.

The west Ramadillas bridge is an older structure that was closed to vehicular traffic at the time of 

the earthquake. The total length is approximately 850 ft (260 m) with 14 spans of about 60.7 ft 

(18.5 m). The superstructure is composed of four parallel steel girders simply supported at the piers

and a reinforced concrete deck. The substructure is composed of two seat-type abutments and

13 intermediate wall-type piers (see figure 77 and figure 78). Steel X-braced cross frames were 

provided at the ends of each span as well as at midspan locations. The girders were anchored to the 

piers by two anchor bolts, one on each side of the bottom flange.

Figure 77. Photo. Overview of west Ramadillas bridge.

Wall-type pier

Steel girders 

and X-bracing

Sand boils
49

Figure 78. Photo. Abutment damage beneath west Ramadillas bridge.

In addition to large inertial forces, significant settlement and lateral flow occurred due to 

liquefaction during the earthquake. As a consequence, one of the wall piers tilted significantly,

and the adjacent span was unseated, as shown in figure 79 and figure 80.

Figure 79. Photo. Pier rotation in collapsed span of west Ramadillas bridge.

Figure 80. Photo. Sand boils near collapsed span in west Ramadillas bridge.

Sand boils

Column rotation Collapsed slab

Steel girders 

and X-bracing

Concrete spalling
50

The largest settlements and lateral flows were observed in the river bank under the south 

abutments of both bridges. Figure 81 through figure 83 show the extent of these effects.

Figure 81. Photo. Settlement at south abutment under east Ramadillas bridge.

Figure 82. Photo. Sliding at south abutment under west Ramadillas bridge.

 

Figure 83. Photo. Block flow toward river due to liquefaction on south bank 

of Ramadillas bridges.

Soil sliding

Soil sliding

Soil surface 

before 

earthquake

Soil sliding

Soil sliding
51

4.3 PERFORMANCE OF SKEWED AND CURVED CONCRETE AND STEEL BRIDGES

4.3.1 Américo Vespucio/Miraflores

At the Miraflores bridge site, two parallel bridges collapsed, as shown in figure 84 through figure 88.

These bridges for southwest-bound and northeast-bound traffic did not have end diaphragms and

appear to have had an S-shaped superelevation. Each three-span precast prestressed girder bridge

had five discontinuous girders and a continuous deck slab. The abutments and two intermediate 

piers were skewed to approximately 20 degrees. Each girder was supported on a neoprene pad 

and a concrete pedestal at both ends. The girders were restrained by a pair of steel stoppers with 

0.4- and 2-inch (10- and 50-mm) gaps for vertical and horizontal motions. Details of the steel 

stoppers are shown in figure 26.

Source: MOP

Figure 84. Photo. Northeast end of Miraflores bridge.

Source: MOP

Figure 85. Photo. Southwest end of Miraflores bridge.

Source: MOP

Figure 86. Photo. Top view of Miraflores bridge from far side

SW abutment

Bottom view

NE abutment

Top view
52

Source: MOP

Figure 87. Photo. Bottom view of Miraflores bridge from near side.

Source: MOP

Figure 88. Photo. Collapsed span of Miraflores bridge.

During the earthquake, one end span of each bridge fell from its seat-type abutment support. The 

other end span of one bridge was partially unseated. After the spans were removed, it was clear that

the entire bridge superstructure rotated clockwise in plan as the acute corner of the bridge moved 

away from its supporting abutment (see figure 89 through figure 91). As a result, the curtain walls

on all abutments suffered damage in the acute corners. The steel stoppers experienced little 

bending deformation, but their anchor bolts were either substantially deformed or sheared off.

Figure 89. Photo. Miraflores bridge after removal of superstructure.

Abutment curtain 

walls damaged

NE abutment
53

Figure 90. Photo. Acute corner of southwest abutment of Miraflores bridge.

Figure 91. Photo. Obtuse corner of southwest abutment of Miraflores bridge.

4.3.2 Américo Vespucio/Lo Echevers

At the Lo Echevers bridge site, two parallel bridges had spans of 91.3, 119.1, and 91.3 ft (27.84, 

36.32, and 27.84 m). As shown in figure 92 and figure 93, the southwest-bound bridge suffered 

minor damage while the northeast-bound bridge collapsed. The bridges did not have end

diaphragms and appeared to have had an S-shaped superelevation. Each three-span precast 

prestressed girder bridge had five discontinuous girders and a continuous deck slab. The 

abutments and two intermediate piers were skewed approximately 33 degrees. Each girder was 

supported on a neoprene pad and a concrete pedestal at both ends. The girders were restrained by 

a pair of steel stoppers with 0.4- and 2-inch (10- and 50-mm) gaps for vertical and horizontal 

motions. Details of the steel stoppers are shown in figure 26.

Source: MOP

Figure 92. Photo. Collapsed northeast-bound Lo Echevers bridge.

Acute corner (damage 

on curtain wall)

Steel Stopper

Obtuse corner (no damage 

on curtain wall)

Concrete Pedestal
54

Source: MOP

Figure 93. Photo. Three unseated spans of northeast-bound Lo Echevers bridge.

During the earthquake, the bridge superstructure rotated clockwise in plan as the acute corners of 

the bridge moved away from their supporting abutments. Some of the anchor bolts fractured, as 

shown in figure 94, and the elastomeric bearings seemed to have been displaced during the 

earthquake, as shown in figure 95. These bearings showed significant lateral bulging and may 

have been overstressed under gravity loads. 

Figure 94. Photo. Failure of a steel stopper on Lo Echevers bridge.

Figure 95. Photo. Displaced elastomeric bearing on Lo Echevers bridge.

Since the two bridges seem identical and experienced similar ground motion, it was unclear from 

visual inspection why one collapsed and the other did not. A numerical analysis of a complete 

model of the bridge may better explain the damage pattern.
55

4.3.3 Avenida Romero Acceso Sur Overpass

The Romero overpass is a two-span bridge with discontinuous precast prestressed girders, a

continuous deck slab, and spans of 96.8 ft (29.5 m) each. Oriented approximately east-west, the 

bridge is similar to the Chada overpass described in section 4.2.2 except that it has a skew angle 

of 31.3 degrees. The bridge has five girders but no diaphragms. Two seismic bars were provided 

between each pair of adjacent girders. As shown in figure 96, the intermediate bent has four

columns that are supported by four drilled shafts with a pile cap between the columns and the 

drilled shaft foundation. Each abutment also rests on four drilled shafts. During the earthquake, 

the superstructure moved toward the west abutment and rotated counterclockwise away from the 

acute corner of the bridge. Both spans were unseated at their abutments, as shown in figure 97, 

which also shows lateral shear failures that occurred in the webs of the exterior girders over the 

bent due to excessive transverse loads. 

Figure 96. Illustration. Intermediate bent details for Romero bridge.

Source: MOP

Figure 97. Photo. Collapse of Romero bridge.

Out-of-plane block 

shear failure

View on 

far side

View on 

near side
56

These observations are supported by the pounding marks on the west abutment, the damage pattern

in the curtain walls of the two abutments, and the orientation of the remaining seismic bars at the 

intermediate bent and two abutments, as shown in figure 98 through figure 100. Figure 101 shows

damage to the back wall and wing wall resulting from the girder impact on the west abutment. 

This damage pattern may have resulted from the combined effects of ground motion characteristics

and the skew configuration. More specifically, the rotational mode of vibration of the bridge

seemed very sensitive to the ground motions, which can be augmented by the skewed abutments. 

In addition, rotational ground motion is another plausible reason for superstructure rotation but 

needs to be verified by examination of the ground motion records.

Figure 98. Photo. Seismic bars at west abutment of Romero bridge.

Figure 99. Photo. Seismic bars at east abutment of Romero bridge.

Figure 100. Photo. Pounding at west abutment of Romero bridge.

Pounding 

marks

West abutment

Drilled shaft

Seismic bars

Curtain wall 

intact

Curtain wall 

damaged

East abutment

Seismic bars

Shear key 

knocked off

Shear key 

intact

Curtain wall damaged Curtain wall intact
57

Figure 101. Photo. Back wall and wing wall damage at west abutment of Romero bridge.

4.3.4 Route 5 Railway Overcrossing at Hospital

The Route 5 railway overcrossing at Hospital comprised three parallel bridges over two railway 

tracks and a local road. The bridges were oriented northwest-southeast. One bridge had a steel 

plate girder superstructure, and the other two had precast prestressed girders. The two concrete 

bridges carried northbound and southbound traffic on Route 5. During the earthquake, the 

southbound bridge collapsed, as shown in figure 102 and figure 103.

Figure 102. Photo. Three bridges (one demolished) at Route 5 overcrossing at Hospital.

Source: MOP

Figure 103. Photo. Collapse of bridge at Route 5 overcrossing at Hospital.

South abutment

Damages

West abutment
58

As shown in figure 102, plastic hinges appear to have developed at the top of the columns of the 

steel plate girder bridge with cross frames and shear keys. The fact that hinges formed at the top 

of the columns implies that strong shaking occurred transversely (i.e., in the direction of the 

skew). This is one of the few bridges for which column damage was observed, but field data is 

minimal because the bridge was demolished for safety reasons soon after the main shock. The 

only bridge to remain standing at this site after the earthquake had diaphragms but no skew.

The collapsed concrete bridge was a two-span structure skewed approximately 40 degrees. The 

superstructure comprised four precast girders but no diaphragms. The intermediate bent comprised

three columns that rested on drilled shafts. At the bent and two abutments, the bridge deck was 

vertically tied to the substructure by seismic bars between pairs of adjacent girders. As shown in 

figure 103, the south span partially fell from the south abutment. The north span was completely 

unseated. The superstructure appeared to have rotated clockwise away from the acute corner, an 

observation confirmed after the bridge deck had been removed (see figure 104 and figure 105).

Figure 104. Photo. Damage at north abutment of Route 5 overcrossing at Hospital.

Pounding 

on backwall

Acute corner curtain 

wall knocked off
59

Figure 105. Photo. Damage at south abutment of Route 5 overcrossing at Hospital.

4.3.5 Quilicura Railway Overcrossing at Avenida Manuel Antonio Matta

The Quilicura railway overcrossing at the Avenida Manuel Antonio Matta, located north of 

Santiago, is a three-span steel bridge with five discontinuous girders, a continuous deck slab, 

and seat-type abutments, as shown in figure 106. The abutments and two intermediate piers are 

skewed approximately 45 degrees. Steel diaphragms are provided at each support, and cross 

frames and stiffeners are evenly spaced along the spans, as shown in figure 107 and figure 108. 

The intermediate piers have five circular columns and a bent cap where the simply supported 

girders are anchored.

Figure 106. Photo. Overview of Quilicura railway overcrossing.

Acute corner curtain 

Obtuse corner wall knocked off

curtain wall intact 
60

Figure 107. Photo. Cross frames on Quilicura railway overcrossing.

Figure 108. Photo. Stiffeners on Quilicura railway overcrossing.

During the earthquake, the superstructure moved both longitudinally and transversely and rotated

away from the acute corner of the bridge. The end span on the east side was unseated at the abutment 

(see figure 109), and the span on the west side was laterally displaced and almost unseated (see 

figure 110). For safety reasons, this span was lifted off the abutment about a day after the main 

shock and was lowered onto a bank of sand placed under the bridge, as shown in figure 111.

Figure 109. Photo. East abutment of Quilicura railway overcrossing.

Curtain wall 

damage

Bridge 

acute corner

Stiffeners

Bent columns

Steel diaphragm

Cross frames
61

Figure 110. Photo. West abutment of Quilicura railway overcrossing.

Figure 111. Photo. West span of Quilicura railway overcrossing lowered onto sandbank.

Due to the transverse displacement and rotation of the superstructure, the curtain walls at the 

acute corners of the bridge sustained pounding damage. In addition, the lateral forces on the 

superstructure damaged the girder anchor bolts and buckled some cross frames, as shown in 

figure 112 and figure 113.

Figure 112. Photo. Damage to anchor bolts on Quilicura railway overcrossing.

Anchor bolt 

failure

Sand

Curtain wall 

damage

Acute 

corner
62

Figure 113. Photo. Damage to cross frame on Quilicura railway overcrossing.

4.4 PERFORMANCE OF STEEL BRIDGES

4.4.1 Tubul Bridge

The Tubul bridge is an eight-span simply supported steel girder bridge oriented north-south with a

span length of approximately 92 ft (28 m). Each span consists of three steel girders with concrete end

diaphragms, as shown in figure 114. Shown in the inset of figure 114 are four seismic bars between

the end diaphragm and the wall pier. The bridge is most likely supported on pile foundations. All 

the spans were unseated at their south end, resulting in the northern-most span punching into the

north abutment and in the buckling of nearby pavement, as shown in figure 115 and figure 116. 

With the exception of span 6, all the spans were in general alignment in the north-south direction. 

Figure 114. Photo. Unseating at southern ends of each span of Tubul bridge.

Figure 115. Photo. Punching of Tubul bridge deck into back wall

Cross frame buckling

N. abutment

Seismic bars End diaphragm

Span 2

Span 8

Span 6

Span 3
63

Figure 116. Photo. Buckling of pavement at north abutment of Tubul bridge.

Local damage to the various piers and spans was inspected by boat. As shown in figure 114, 

span 6 lost support at its north end. Local residents confirmed that this span was like the others 

immediately after the main shock, but the north end collapsed during an aftershock during the 

first week, probably due to rigid body rotation of the wall pier. This difference in span 6 may 

be attributable to the fact that the span rotated from the north to the northeast, creating load 

concentration on the north support. Figure 117 shows watermarks on one of the wall piers. The 

concrete below the high watermark showed signs of deterioration, probably due to wet and dry 

cycles, particularly in the mid-height of the wet area at the average watermark.

Figure 117. Photo. Span 6 of Tubul bridge, unseated at north end during an aftershock.

In general, the bottom flange of the girders buckled, as shown in figure 115 and figure 118, and 

the tops of the wall piers were damaged locally, as shown in figure 118 and figure 119. All of the 

piers were tilted toward the north abutment and were significantly damaged either by flexural 

effects at the top of the footing due to the maximum moment as the bridge deck pushed against 

the wall during the earthquake or by shear effects around the average watermark where the concrete 

was severely deteriorated (see figure 119 and figure 120). Concrete spalling associated with both 

a shear crack and a vertical crack on the wall pier are shown in figure 121. The pile cap holding 

the dropped span was cracked, as shown in figure 122, probably due to the impact of the falling span.

Span 6

High watermark

Average watermark

Top of footing
64

Figure 118. Photo. Bottom flange buckling and concrete spalling on Tubul bridge.

Figure 119. Photo. Shear crack in Tubul bridge.

Figure 120. Photo. Tilting wall pier on Tubul bridge.
65

Figure 121. Photo. Spalling due to shear action and vertical crack in Tubul bridge.

Figure 122. Photo. Crack in footing of Tubul bridge.

4.4.2 Cardenal Raúl Silva Henríquez Bridge

The Cardenal Raúl Silva Henríquez bridge, built in 2002, is a 22-span steel girder structure that 

crosses the Maule River in the northeast-southwest direction near Constitución. It is supported by 

2 seat-type abutments and 21 intermediate bents (see figure 123). The first five bents from the 

northeast abutment are supported on two reinforced concrete columns and drilled shafts. The next

six bents are supported on three reinforced concrete columns and drilled shafts. The following 

eight bents are steel pile bents with three legs (one vertical and two inclined) and horizontal struts

and diagonal braces interconnecting the legs in each bent. The last two bents are supported on 

three reinforced concrete columns that rest on footings. The superstructure comprises two

continuous segments of 11 spans each. There are fixed joints at each abutment and an expansion 

joint at the middle of the bridge over bent 11. Note that girders with fixed bearings at an abutment

seem to defeat the purpose of having expansion joints at the abutment. At each abutment, the 

bottom flanges of the three girders are welded to masonry plates embedded in the abutment seat. 

An elastomeric pad is provided over each intermediate bent to permit longitudinal movement. 

Transverse and vertical movements are restrained by steel stoppers, as shown in figure 124. 
66

Figure 123. Photo. Cardenal Raúl Silva Henríquez bridge looking south.

Figure 124. Photo. Elastomeric pad and stopper over bent 

on Cardenal Raúl Silva Henríquez bridge.

During the earthquake, the northeast portion of the bridge moved from west to east in the transverse

direction. All steel stoppers were deformed, and girders were displaced from their elastomeric 

pads. The transverse diaphragms at the intermediate bents and abutments buckled, as shown in 

figure 125 and figure 126. At the northeast abutment, the webs and bottom flanges were fractured

in all three steel girders, and both bearing stiffener and web buckling occurred, as shown in 

figure 127. The exterior girder also experienced web and flange bending about its weak axis, as 

shown in figure 128. Web buckling was observed in the longitudinal direction in the same girder. 

This type of damage is evidence of large longitudinal loads in the superstructure being attracted 

to the fixed bearing at this abutment and that these loads changed sign from tension to compression

several times during the long-duration event. The weld between the girders and the masonry 

plate did not fail. 

NE

SW
67

Figure 125. Photo. Girder offset and cross frame buckling 

on Cardenal Raúl Silva Henríquez bridge.

Figure 126. Photo. Cross frame buckling on Cardenal Raúl Silva Henríquez bridge.

Buckling o f 

end diaphragm

Buckling of end 

diaphragm
68

Figure 127. Photo. Girder damage at northeast abutment 

of Cardenal Raúl Silva Henríquez bridge.

Figure 128. Photo. Temporary repair at northeast abutment 

of Cardenal Raúl Silva Henríquez bridge

To maintain traffic on the bridge, a temporary cross frame was installed at the northeast abutment, 

as shown in figure 128. Due to severe damage in the girder webs and flanges, a steel beam was 

used to transfer loads from the girder flanges to the abutment seat, bypassing the webs.

As previously noted, nine spans of the southwest portion of the bridge are supported on tall steel pile

bents, as shown in figure 129. The legs of these bents are fabricated from 4.9-ft (1.5-m)-diameter 

tubes with a wall thickness of 0.55 inches (14 mm). Their overall heights from pile tip to soffit 

range from 146 to 169 ft (44.5 to 51.5 m). The maximum clear height above water is about 82 ft 

(25 m). This portion of the structure appeared to have performed well during the earthquake but 

suffered local damage at several locations due to tsunamis. Four major waves were recorded nearby,

with arrival times ranging from 4 to 5.5 h after the earthquake. Wave heights ranged from 22.6 to 

36.7 ft (6.9 to 11.2 m), passing under the superstructure. Nevertheless, the waves eroded river 

sands and gravels from around the piles to a depth of about 15 ft (4.5 m), and waveborne debris 

punctured a hole in the wall of one of the legs of one of the bents (see figure 130 and figure 131).
69

Source: MOP

Figure 129. Photo. Spans and pile bents of southwest portion 

of Cardenal Raúl Silva Henríquez bridge.

Source: MOP

Figure 130. Photo. Erosion of alluvial material around bent legs 

of Cardenal Raúl Silva Henríquez bridge.

Source: MOP

Figure 131. Photo. Debris-impact hole in bent leg of Cardenal Raúl Silva Henríquez bridge.
70

Although the details are similar, the damage to the girders at the southwest abutment was not as 

severe as at the northeast abutment and was limited to the buckling of the bearing stiffeners. The 

reason for this difference is probably that the fillet welds connecting the girder flanges and bearing

stiffeners to the masonry plates failed during the earthquake and large longitudinal loads were not

attracted to this abutment (see figure 132 through figure 134).

Source: MOP

Figure 132. Photo. Fillet weld fractures at the girder-to-abutment connection on the 

Cardenal Raúl Silva Henríquez bridge.

Source: MOP

Figure 133. Photo. Close-up of girder end on Cardenal Raúl Silva Henríquez bridge.

Source: MOP

Figure 134. Photo. Close-up of transverse stiffeners on Cardenal Raúl Silva Henríquez bridge.
71

4.4.3 Biobío River Bridge

The first bridge over the Biobío River in Concepción was built in 1925 and consisted of 108 simply

supported spans of approximately 49 ft (15 m) each. Each span had a timber superstructure 

(beams and deck) and rested on concrete piers that were supported on battered timber piles. After 

40 years of service, the Chilean Highway Authority decided to replace the bridge with a steel-girder

structure of 90 spans (approximately 49 ft (15 m) each), which were also supported on wall piers 

and timber pile foundations. During the 1960 earthquake, the five spans closest to Concepción 

collapsed. In 1965, an additional three spans collapsed due to scour effects. As a result, most of 

the pile foundations were replaced in the mid-1960s with steel pipe piles and concrete cap beams.

In 1994, Japanese consultants evaluated the structural integrity of the bridge and investigated 

various retrofitting options. It was concluded that the bridge should be closed to traffic and was 

too expensive to retrofit due to the deterioration of many structural components. Nevertheless, 

local authorities decided to keep the bridge open for limited traffic such as pedestrians and 

bicycles. After additional studies, the bridge was completely closed in 2002, and all traffic was 

detoured to the Llacolen bridge, which had been built in 2000. In 2009, MOP contracted the firm

Q&R Ingeniería to conduct a preliminary study for the replacement of the Biobío bridge.

The historic multispan bridge over the Biobío River is oriented approximately east-west. As seen 

in figure 135 and figure 136, the bridge superstructure consists of multiple simply supported spans

with four steel girders and a concrete deck. Intermediate and end X-braced cross frames were

provided in each span. Each end of a steel girder was restrained by two anchor bolts through the 

bottom flange, one on either side of the web. As shown in figure 137, lateral spreading due to 

liquefaction was observed near the north end of the bridge approaches. Many spans moved

longitudinally and fell from their seats after shear failure of the anchor bolts. Some piers also 

collapsed in the longitudinal direction, as shown in figure 138, possibly due to liquefactioninduced vertical settlement in combination with longitudinal seismic forces. These movements

may have caused the multiple cracks visible in the pile caps (see figure 139).

Figure 135. Photo. Collapsed spans in Biobío River bridge.
72

Figure 136. Photo. Unseated spans in Biobío River bridge.

Figure 137. Photo. Lateral spreading near east end of Biobío River bridge.

Figure 138. Photo. Collapsed spans and piers of Biobío River bridge.

Figure 139. Photo. Cracks in pile caps of Biobío River bridge.
73

4.4.4 Pichibudis Bridge

The Pichibudis bridge is located on a coastal road (Route J-60) in Punta Duao near Iloca. The 

bridge crosses a stream 980 ft (300 m) before the stream flows into the sea (see figure 140). 

© Google, DigitalGlobe, Inav/Geosistemas SRL, and DMapas

Figure 140. Photo. Satellite image of Pichibudis bridge.

The single-span bridge has a total length of approximately 98 ft (30 m). The superstructure is 

composed of two parallel 62-inch (1,600-mm)-tall plate girders and a reinforced concrete deck 

slab. The girders are detailed with cross frames and stiffeners evenly spaced along the span, as 

shown in figure 141 and figure 142. The supports are seat-type abutments, reinforced concrete 

diaphragms, and concentric braces fabricated from steel reinforcing bars. The bottom flange of 

each girder sits on rubber pads at each abutment. 

Figure 141. Photo. Cross frames in Pichibudis bridge.
74

Figure 142. Photo. Web stiffeners in Pichibudis bridge.

The bridge was overtopped by one or more tsunami waves, and the deck slab suffered a permanent 

lateral displacement of about 7 inches (180 mm) at the south abutment (see figure 143 and

figure 144). At this abutment, the seaward girder rotated about the bottom flange and suffered

flexural and torsional deformations about the weak axis of the section. 

1 cm = 0.39 inches

Figure 143. Photo. Offset in girder top flange of south abutment of Pichibudis bridge.

Figure 144. Photo. Offset in handrail on south abutment of Pichibudis bridge.

18 cm

Lateral Offset

Girder 

rotation

18 cm

To the sea
75

At the same location, damage was also found in the concrete diaphragm and bracing system, with

tension rupture in one bar and compression buckling in the other one (see figure 145 and figure 146).

Figure 145. Photo. X-braces damage in south abutment of Pichibudis bridge.

Figure 146. Photo. Diaphragm crushing in south abutment of Pichibudis bridge.

According to local residents, the tsunami was characterized by three waves starting approximately

10 min after the main shock. Since the clear height of the bridge from the water level is about 4.9 ft

(1.5 m), the tsunami waves passed over the bridge, inducing large lateral forces on the girders. In 

addition, a visual inspection of the connecting elements at the south abutment revealed corrosion

in the girders, cross braces, and vertical restrainers in the reinforced concrete diaphragms (see 

figure 147). Due to the corrosion, the capacity of the structural components of the south abutment

was less than at the north abutment, making this abutment more likely to suffer damage.

Compression 

buckling

Tension 

rupture

Girder 

Diaphragm damage

damage
76

 

Figure 147. Photo. Damage on south abutment due to corrosion in Pichibudis bridge.

4.4.5 El Bar Bridge

The El Bar bridge is located on Route 160 in Arauco and crosses a small lake in the province of 

Carampangue, as shown in figure 148. 

© Google, GeoEye, and DMapas

Figure 148. Photo. Satellite image of El Bar bridge.

Originally designed as a single-span bridge with a total length of approximately 98 ft (30 m), an 

intermediate support had recently been added (see figure 149). Apparently, this bridge suffered 

structural damage during flooding in 2006 and was retrofitted accordingly. In addition to the 

midspan support frame, two new plate girders were added parallel and adjacent to the existing 

girders. The new girders were connected to the existing girders using welded steel angles, as 

shown in figure 150. The connection between the new beams and the deck slab was made using 

steel angles and bolts anchored with epoxy, as shown in figure 151.

Girder corrosion

Vertical restrainer 

fractured

5 cm
77

Figure 149. Photo. Side view of El Bar bridge.

Figure 150. Photo. Retrofit girder connections on El Bar bridge.

Figure 151. Photo. Retrofit deck connections on El Bar bridge.

Earthquake damage at the south abutment included a lateral permanent deformation of 9.8 inches 

(250 mm), girders dislodged from their neoprene pads, and damage to the concrete back wall (see

figure 152 and figure 153). When the bridge was retrofitted after the 2006 floods, a rebar X-brace

system and a new concrete diaphragm were added at the north abutment. The diaphragm was 

Deck connectors

Bearing stiffener

Old girder

New girder

Connector

New girder

Midspan structure
78

anchored to the abutment seat using vertical restraining bars (seismic bars). Earthquake damage 

observed at this support included large deformations in the rebar of the X-brace system (tension 

yielding and compression buckling), lateral deformation of the seismic restrainers, and concrete 

crushing of the original abutment back wall (see figure 154 and figure 155). 

1 cm = 0.39 inches

Figure 152. Photo. Lateral deformation at south abutment of El Bar bridge.

Figure 153. Photo. Damage at south abutment of El Bar bridge.

Figure 154. Photo. Damage to rebar X-brace system on north abutment of El Bar bridge.

Rubber 

pad

Girder 

sliding off

Back wall 

damage

25 cm

25 cm
79

Figure 155. Photo. Damage to north abutment back wall of El Bar bridge.

4.4.6 Itata River Bridge

Built in 1980s, the Itata River bridge crosses the river near Coelemu (see figure 156). The total 

length of the bridge is approximately 2,160 ft (660 m) made up of 22 98-ft (30-m) spans. The 

superstructure is composed of two parallel steel plate girders simply supported between piers 

and a reinforced concrete deck slab. The substructures include two seat-type abutments and 

21 intermediate 39-ft (12-m)-tall wall-type piers (see figure 157). The two girders have equally 

spaced web stiffeners and are interconnected by inverted chevron-type cross frames. Reinforced 

concrete diaphragms are provided at the abutments and over each pier.

© Google, DMapas, DigitalGlobe, Inav/Geosistemas SRL

Figure 156. Photo. Satellite image of Itata River bridge.

Damage on the back wall
80

 

Figure 157. Photo. Side view and superstructure details for Itata River bridge.

Each girder sits on a reinforced concrete bracket at the top of each pier (see figure 158). These 

brackets were designed to transfer the axial and lateral forces to the piers, which are founded on 

rectangular pile caps.

During the earthquake, the entire structure moved both longitudinally and transversely, with larger

displacements occurring in the transverse direction. As a consequence, many of the concrete brackets

suffered damage at their connections to the piers. The damage was characterized by shear cracks, 

concrete spalling, and bar buckling, as shown in figure 159. At the time of inspection, a temporary

steel structure was under construction to retrofit the brackets and allow traffic on the bridge. 

 

Figure 158. Photo. Girder supports at each pier of Itata River bridge.

RC Brackets

RC Diaphragm

Cross frames

Wall-type pier

W. Abutment

E. Abutment

Wall-type pier. 
81

 

Figure 159. Photo. Damage to girder supports of Itata River bridge.

4.4.7 Pedestrian Bridge Over Route 5

Several pedestrian bridges over Route 5 collapsed during the earthquake. All appeared to have 

similar details and to have failed in a similar manner. The collapse of the pedestrian bridge at 

KM 13.8 of Route 5 North is described in this section. The bridge consisted of two approach 

ramps on each side of the highway and a two-span overpass oriented approximately east-west. 

As shown in figure 160, the two spans fell from their supports during the earthquake and were 

subsequently moved to the sides of the highway. Each 70.1-ft (21.4-m)-long span consisted of a 

concrete deck supported by two steel plate girders and four uniformly spaced cross frames. The 

end of each girder rested on a neoprene pad, as shown in figure 161. Each cross frame was composed

of X-bracing, a top chord, and a bottom chord, as shown in figure 162. The main span was 

anchored at each end to a hammerhead pier by two 1-inch (25-mm)-diameter bolts through the 

flange of the bottom chord. Failure occurred when the anchor bolts pulled out of the bottom chord

and there was insufficient seat width to prevent the girders from collapsing (see figure 163).

Figure 160. Photo. Damaged Route 5 pedestrian bridge.

Concrete spalling

Shear cracks

Bar buckling
82

Figure 161. Photo. Steel girder on neoprene pad at abutment of Route 5 pedestrian bridge.

Figure 162. Photo. End cross frame near collapsed span of Route 5 pedestrian bridge.

Figure 163. Photo. Anchor bolt failure (shear and pull out) on Route 5 pedestrian bridge.

4.5 PERFORMANCE OF OTHER BRIDGE TYPES

4.5.1 Claro River Bridges

Three arch bridges cross the Claro River just south of Curicó. Two of these are shown in figure 164:

a reinforced concrete highway arch that carries northbound and southbound lanes of Route 5 and 

a set of unreinforced masonry arches. Built in 1870, the masonry bridge is a historic structure. It

has survived numerous strong earthquakes in its lifetime, including the 1960 and 1985 events. 

The reinforced concrete arch suffered only minor damage during the 2010 earthquake, but the 

masonry bridge collapsed completely (see figure 165).

Cross frame

Broken anchor bolt
83

Figure 164. Illustration. Highway bridges crossing Claro River before earthquake.

Figure 165. Photo. Collapse of masonry arch bridge at Claro River.

The masonry bridge consisted of a set of seven back-to-back arches oriented approximately 

northeast-southwest. Two piers survived the earthquake, and all foundations appeared to be 

intact, as shown in figure 165 and figure 166. 

 

Figure 166. Photo. Close-up view of collapsed bridge at Claro River.

Unreinforced masonry 

arch bridge

Reinforced concrete 

arch bridge

Steel arch 

railway bridge

Pier 1 Pier 2 Pier 3 Pier 4 Pier 5 Pier 6

Southwest (SW)

abutment

Northeast (NE)

abutment

Southwest (SW)

abutment

Pier 1

Pier 2

Pier 3

Pier 4

Pier 5

Pier 6
84

Figure 167 through figure 169 show details of various components of the bridge. The deck slab

was approximately 8 to 10 inches (203 to 254 mm) thick, and the arch crown was approximately 

4 ft (1.2 m) thick. The ribs and piers were constructed from solid brick with some stone fill above

the brick ribs, as shown in figure 167 and figure 168, and faced with stone blocks or plastered 

with cement to give the appearance of stonework. Close inspection of the piers found some cracks

between the bricks and stone facades, possibly due to shaking, but, in general, the masonry was 

in excellent condition. The foundations of the bridge were most likely made of stone blocks and 

seemed to have experienced little or no damage during the earthquake (see figure 169). 

1 m = 3.28 ft

Figure 167. Photo. Deck slab and crown of arch on Claro River bridge.

Figure 168. Photo. Pier of Claro River bridge.

~ 4 ft Brick arch

Stone fillers
85

Figure 169. Photo. Foundation of Claro River bridge.

As shown in figure 166, the collapsed spans and other debris, including arches and piers, fell 

directly into the river below the bridge and were not thrown sideways, upstream, or downstream. 

Therefore, one scenario is that the collapse was caused by the failure of a pier with deteriorated 

masonry during strong shaking along the axis of the bridge. Another scenario is that the elevation 

difference between the end support of the arch bridge and its nearby pier caused stress concentration

in two end spans along the axis of the bridge, resulting in cracks in the ribs and falling of various 

pieces. The collapse of two end spans reduced the inward thrust of each intermediate arch span 

and triggered progressive collapse. Since no deterioration of the masonry was observed in the 

rubble in the river bed, the second scenario is most likely. 

Based on field observations, the progressive collapse of the entire masonry bridge could have 

occurred in four steps, as shown in figure 170. The figure shows the pre-earthquake condition of 

the bridge, indicating the horizontal inward thrust of each arch under gravity loads. Due to the 

difference in support conditions in the two end spans, the bending moment distribution along the 

axis of the arch bridge was quite uneven under the earthquake load. The maximum moment most 

likely occurred in the end spans. As a result, each end arch was subjected to cracking, leading to 

loss of various pieces (see step 1 of figure 170). As the end spans collapsed, the sudden loss of 

thrust for the adjacent spans caused significant force redistribution in a short time, resulting in 

cracking and collapse of the spans under gravity or earthquake loads, as illustrated in step 2.

Similarly, the next spans collapsed, as shown in step 3.

Next, the center span experienced force redistribution due to the collapse of its adjacent spans. 

However, the change in horizontal thrust for the center span was likely smaller than that in the

other intermediate spans. In addition, the earthquake excitation was likely over by that time. 

Therefore, the arch collapsed mainly due to its own weight while the two piers supporting the 

center span remained, as shown in step 4. When compared with figure 165, this collapse matches 

the field observations.
86

Pre-earthquake condition 

Step 1

Step 2

Step 3

Step 4

Figure 170. Illustration. Scenario for the progressive collapse of Claro River bridge.

4b 4a

3b 3a

2b 2a

1b 1a
87

4.5.2 Chepe Railroad Bridge Over Biobío River

The Chepe railroad truss bridge over the Biobío River was built in 1889 and retrofitted in 2005. 

Oriented northeast-southwest, the bridge has two continuous through-truss spans at the north end 

and multiple truss spans for the river crossing (see figure 171 through figure 173). The throughtruss structure is supported on the north abutment and two intermediate bents. Each bent comprises

two concrete-filled steel tubes that are 72 inches (1,807 mm) in diameter, as shown in figure 174. 

All the river spans are supported on steel towers. Each tower comprises six steel circular piles 

that are cross-braced for lateral strength and stiffness (see figure 172 and figure 175). A view 

underneath the through-truss spans and a sectional view of the remaining spans are shown in 

figure 174 and figure 175.

Figure 171. Photo. North end of Chepe railroad bridge over roadway.

Figure 172. Photo. River spans of Chepe railroad bridge.

Figure 173. Photo. South end of Chepe railroad bridge.

Figure 174. Photo. Bottom view and cross section on north end of Chepe railroad bridge.

Concrete-filled 

steel tube 

Bent with six

 steel piles
88

Figure 175. Photo. Section on south end of Chepe railroad bridge.

At the northeast end, the bridge spans over Route O-60, which runs in a trench between two retaining

walls, as shown in figure 176. The roadway slab is supported by micropiles that are approximately

20 ft (6.1 m) long. During the earthquake, the steel tube bent behind the river-side retaining wall 

settled approximately 12 inches (300 mm) and moved laterally toward the river for approximately

26 inches (660 m). It also tilted toward the river about 5 degrees, as shown in figure 177. In the 

vicinity of the steel tube bent, the soils behind the wall settled approximately 51 inches (1,300 mm)

relative to the steel bent, as shown in figure 177. The retaining wall settled over 12 inches (300 mm),

as shown in figure 178, and moved laterally toward the river for approximately 26 inches (660 mm),

as shown in figure 176. As a result, the seal between the roadway and toe of the retaining wall 

sustained damage, creating the possibility for water to flood the roadway as the river level rises 

during the winter. The movement of the steel tubes and retaining wall was due to liquefactioninduced ground spreading.

Steel Caisson

Micro Piles

Wall O-60 Roadway Slab

Seal

Ground Spreading

Lateral movement

and settlement of

approach pier

Lateral movement

and settlement of

Retaining wall

Figure 176. Illustration. Settlement and lateral movement of wall and bent due to ground 

spreading under Chepe railroad bridge.

Bent with six 

steel piles

Cracks
89

Figure 177. Photo. Settlement and tilting of bent on Chepe railroad bridge.

Figure 178. Photo. Settlement behind retaining wall on Chepe railroad bridge.

At the south end of the bridge, several bracing rods were broken, and at least two steel piles were 

fractured, possibly due to ground spreading, as shown in figure 179 and figure 180. A temporary 

support comprising a stack of timber ties had been erected near the damaged bent, as shown in

figure 173.

Relative wall 

settlement ~12 ̋ 

Lateral 

movement

51 ̋ 

settlement

26 ̋ 

movement

Tilted 

toward river

Centerline 

of the bent
90

Figure 179. Photo. Ruptured cross bracing at south end of Chepe railroad bridge.

Figure 180. Photo. Fractured steel piles at south end of Chepe railroad bridge.

4.5.3 Maipú River Bridges

Three bridges cross the Maipú River just south of Santiago: a railway bridge, a highway bridge 

(Route 5), and a local access road bridge (see figure 181). 

Figure 181. Photo. Three parallel bridges crossing Maipú River south of Santiago.

The railway bridge seemed to be undamaged by the earthquake. Likewise, the Route 5 bridge 

sustained only minor damage despite being built in 1970. However, the Route 5 bridge is susceptible 

to scour, and the foundations were being retrofitted at the time of the reconnaissance visit. The local

access road bridge was severely damaged in the earthquake. 
91

Built in 1960s, the Maipú River bridge has four continuous segments of 3, 3, 3 and 4 spans, for a 

total of 13 spans. The segments are separated by expansion joints. Each reinforced concrete span is 

composed of three girders with end haunches, six intermediate diaphragms, and two end diaphragms. 

The superstructure is supported by A-shaped reinforced concrete piers, as shown in figure 181.

Oriented approximately north-south, the bridge was used to carry trucks loaded with gravel dredged

from the river. Field observations indicated the existence of shear cracks near the ends of many 

of the girders, probably caused by heavy truck loading (see figure 182 and figure 183). Some of 

these cracks appeared to have been repaired prior to the earthquake. 

Figure 182. Photo. Crushing of haunches at pier 3 of Maipú River bridge.

Figure 183. Photo. Shear cracks at span 1 of Maipú River bridge.

During the earthquake, the haunches of three girders over pier 3 were severely crushed, as shown

in figure 182. Some concrete diaphragms, such as those at piers 2 and 3, were also cracked 

significantly, as shown in figure 184. After the earthquake, two temporary frames were installed 

underneath span 1, and lightweight traffic was permitted on the bridge. Most likely, the crushing 

of the girder haunches was mainly due to a combined effect of seismic-induced longitudinal shear

force and vertical load, as shown in figure 184.

Figure 184. Photo. Longitudinal and vertical seismic forces in haunches at pier 3 

of Maipú River bridge.

93

CHAPTER 5. GEOTECHNICAL PERFORMANCE 

OF TRANSPORTATION INFRASTRUCTURE

5.1 OVERVIEW

Geotechnical issues influencing the degree of damage to the transportation infrastructure include 

amplification of ground motions due to the presence of soft soil or basin effects, excessive lateral 

foundation flexibility due to soft or loose soil conditions, settlement, lateral ground movement, and

ground failure. Settlement, lateral ground movement, and ground failure appeared to be primarily the 

result of soil weakening due to liquefaction but, in some cases, were due to seismically generated 

inertial forces acting on the soil mass for existing soil slopes that may have been marginally stable. 

A significant portion of this chapter is devoted to the effect liquefaction can have on transportation

infrastructure. Liquefaction is the significant loss of soil strength and stiffness resulting from the 

generation of excess pore-water pressure in saturated, predominantly cohesionless soils. Kramer 

provides a detailed description of liquefaction, including types of liquefaction phenomena, 

evaluation of liquefaction susceptibility, and effects of liquefaction.(12)

As the pore-water pressure in the soil builds up during liquefaction, loss of soil strength occurs. 

Then, settlement of the liquefied soil occurs as the excess pore-water pressures dissipate after the 

earthquake. The potential effects of strength loss and settlement include the following: 

Slope failure, flow failure, or lateral spreading. The strength loss associated with porewater pressure buildup can lead to slope instability. If the soil liquefies, the soil strength 

reduces to a very low residual value, and the slope can become unstable. The lateral loads 

due to soil movement created by the unstable soil mass can cause bridge substructure 

distortion and unacceptable deformations and moments in the superstructure.

Reduced foundation bearing resistance. The residual strength of liquefied soil is often 

a fraction of non-liquefied strength. This loss in strength can result in large displacements 

or foundation bearing failure.

Reduced soil stiffness and loss of lateral support for deep foundations. This loss in 

soil strength can change the lateral response characteristics of piles and shafts under 

lateral load, making them much more flexible. This increases deformations and forces in 

the superstructure. 

Vertical ground settlement as excess pore-water pressures induced by liquefaction 

dissipate, resulting in downdrag loads on and a loss of vertical support for deep 

foundations. When liquefaction-induced downdrag occurs, it can cause significant 

foundation settlement, depending on the resistance of the soil at the tips of the 

foundation elements.(13)

In this chapter, these geotechnical issues are addressed for transportation facilities including 

bridges, pier structures, retaining walls, embankments, and roadway cuts. 
94

5.2 GEOTECHNICAL OBSERVATIONS REGARDING BRIDGE AND PIER 

STRUCTURES

Of the 32 sites visited, 15 exhibited signs of liquefaction. Bridge performance tended to be 

poorer when liquefaction occurred than when it did not. The primary cause of transportation 

infrastructure performance problems resulting from liquefaction is the movement of the ground 

laterally and vertically in the vicinity of those facilities. However, liquefaction is not the only 

cause of ground movement during earthquakes. Relatively weak, sensitive clays can also lose 

strength during earthquakes, resulting in potentially large lateral ground movement that can have 

an impact on structures, walls, and roadway fills. Severe ground movements due to the presence 

of weak, sensitive clays were observed at 4 of the 32 sites (i.e., liquefaction was not the cause of 

the movements). There are still some cases of severe bridge damage or collapse among the 13 sites 

where neither liquefaction nor ground movement due to weak, sensitive clays appeared to have 

occurred. This damage was primarily due to structural weaknesses; though in many of those 

cases, performance problems may have worsened due to the presence of soft or loose soils.

Overall, the breakdown of observed geotechnical problems is as follows:

Liquefaction: 15 sites (sites 15–22, 24–28, 30, and 31).

Ground movement due to weak, sensitive clays: 4 sites (sites 7, 8, 10, and 12).

Ground motion amplification likely but no ground movement due to liquefaction or 

weak, sensitive clays: 9 sites (sites 1–6, 11, 23, and 32).

The breakdown of the types of geotechnical issues and problems affecting the bridges visited is 

as follows:

Excessive lateral dynamic foundation movements due to the presence of loose or soft 

soil resulting in structural damage or collapse (without liquefaction): Evidence was 

directly observable at 1 bridge site (site 8) but probably occurred at other bridge sites visited.

Abutment fill settlement or lateral ground movement due to the presence of weak, 

sensitive clays: In all cases, it appeared that the lateral soil movement did not significantly

impact the bridge (i.e., the bridge abutments did not move); 4 sites (sites 7, 8, 10, and 12).

Abutment fill settlement or lateral ground movement due to liquefaction, but with 

minimal liquefaction-induced foundation settlement or ground failure-induced 

foundation lateral displacement and bridge damage: 8 bridge sites (sites 15, 16, 17 

(southwest end), 18 (southwest end), 20, 21, 28, and 30). 

Bridge pier settlement or lateral movement due to liquefaction and ground failure 

resulting in significant bridge damage or collapse: 1 pier (wharf) and 10 bridge sites

(sites 17 (northeast end), 18 (northeast end), 19, 21, 22, 24–27, and 31).

Other geologic effects potentially impacting bridges, including regional ground uplift

or subsidence and tsunami damage: 5 bridge sites (sites 20–22 for regional uplift and

possibly others to a lesser degree, and sites 14 and 16 for tsunami).
95

5.2.1 Non-Liquefaction Geotechnical Issues Observed for Bridges

Bridge damage was generally more common in areas with poorer soils, resulting in greater 

amplification of ground motions. Bridges damaged in the Santiago area appeared to follow this 

pattern, with the heaviest damage observed in the western sector. As discussed in chapter 2, peak 

accelerations in the ground motions from the eastern side of Santiago were only one-third to onehalf the peak accelerations measured in the western sector, based on the preliminary ground 

motion data available. This observation is similar to those summarized in GEER, not only for the 

Santiago area but also for other locations.(2) Depending on the variability of the soil profile below 

each bridge, the amount of amplification may have varied from bridge to bridge, even within a 

given bridge site. Furthermore, the extent of the damage depended on the bridge geometry, 

details used, and design code followed.

5.2.1.1 Bridge Sites 1–6, North and West Santiago

The bridges located on or in the vicinity of the northwestern portion of the Américo Vespucio, 

the beltway around the Santiago metro area, are examples of locations where ground motion 

amplification due to the presence of softer or looser soils was the primary geotechnical issue

affecting the damage observed (see appendix A for boring logs that provide examples of subsurface 

conditions in the area). These include bridges at sites 1–6. Figure 92 and figure 93 show examples

in which the superstructure was damaged or collapsed while the substructure was not significantly

damaged. In these cases, weaknesses in the superstructure design (lack of shear restraint, inadequate 

bridge seat width, etc.) helped reduce dynamic stresses and damage in the substructure as the 

superstructure lost its connection to the substructure during the vigorous shaking.

The bridge design and details were similar for bridge sites 1–3. All of these bridges were likely 

supported on 4.9-ft (1.5-m)-diameter drilled shafts, with the exception of the footing-supported 

Lo Echevers bridges (see appendix A). Why some of these bridges suffered severe damage and 

others did not is unknown at this point, but it is possible that differences in soil conditions beneath

these bridges may have resulted in different degrees of ground motion amplification and therefore

differences in the observed bridge performance. That is, the stronger the ground motions at a site, 

the more likely that bridge damage occurred.

5.2.1.2 Bridge Sites 7–10, South Santiago

At the south end of the Santiago area, soft or loose to medium dense soils were present based on 

the available geologic maps and boring logs (see chapter 2 and appendix A). The boring log 

obtained at sites 7, 8, and 10 illustrate that the upper 38 to 58 ft (11 to 16 m) of soil consists of 

soft to stiff clays or loose to medium dense sandy silts and silty sands (AASHTO and NEHRP 

Site Class D or E), which would provide maximum amplification of ground motions, based on 

the AASHTO guidelines for seismic load and resistance factor design (LRFD).

(14) However, this 

area is on the margin of a region with better soil conditions (granular medium-dense to dense soils),

and therefore, subsurface conditions may be more varied, causing ground motion amplification 

and bridge performance to be more varied. In general, this applies to sites 7–12. For all of these 

bridges, the damage observed was primarily in the superstructure, with no apparent substructure 

damage. The foundations for the bridges at sites 7, 8, and 10 consisted of four 4.9-ft (1.5-m)-diameter

drilled shafts extending into the dense soil (about 36 to 98 ft (11 to 30 m) below roadway grade),
96

based on as-built plans produced in 2000 and 2001 (see appendix A). Figure 185 shows an 

example in which the substructure was left intact with little or no damage. However, one of the 

foundations supporting one of the bridge abutments appears to be less vertical than would be 

expected, though no specific evidence of foundation damage (e.g., cracking) was observed. 

However, in one of the bridges (site 8, Avenida Chada overpass), dynamic lateral movement of 

the center pier shaft foundation had occurred (on the order of about 7.8 inches (200 mm)) and 

lateral foundation movement likely contributed to bridge superstructure damage (see figure 186). 

Figure 187 shows additional evidence that may point to the transverse rocking that occurred at 

the Avenida Chada overpass. The pavement crack shown in the figure is directly below the 

centerline of the bridge. While it is possible that the foundations supporting the bridges at all or 

most of the bridges at sites 1–10 exhibited significant lateral movement due to the presence of 

soft soils, it was not directly observable, except at site 8.

Figure 185. Photo. Damage at Romero bridge and shaft foundation supporting abutment

(collapsed superstructure removed).

Damaged shear walls Fill settlement
97

Figure 186. Photo. Interior pier of Chada structure with 5.8- to 7.8-inch (150- to 200-mm) gap 

in soil transverse to bridge centerline.

Figure 187. Photo. East abutment of Chada structure showing large pavement crack 

directly below bridge centerline.

Another problem that occurred at these bridge sites was settlement of the approach fills, where 

total approach fill settlement was on the order of 3 ft (1.0 m). However, for bridge sites 7 and 8,

this settlement seems to be entirely from within the fill due to the fill slumping during shaking.
98

The fill slumping did not appear to be caused by liquefaction. It is possible that during shaking,

the base of the fill slid away from the embankment centerline due to some type of block failure, 

possibly at the boundary with relatively soft clay soils, allowing the embankment to slump.

Figure 188 is a diagram of fill slumping, and figure 189 shows one of the actual abutment fills.

The type of fill material used here and how well it was compacted is unknown.

Figure 188. Illustration. Fill slumping observed for bridge sites 7 and 8.

Figure 189. Photo. Fill slumping observed at Romero bridge.

Significant roadway fill settlement and lateral movement also occurred at bridge site 10 (see 

figure 190). However, the difference between this site and sites 7 and 8 is that at site 10, it appears 

the soil beneath the roadway embankment failed, possibly in bearing, though the failure appears 

to be fairly shallow. Lateral movements of over 7 ft (2 m) appear to have occurred, considering 

the distortion in the sign bridge, and uplift of the soil surface beyond the toes of the approach fill 

also appears to have occurred (see figure 191). Lateral movement of the fill and soil below the fill

at site 10 is rather complex and will require further investigation to determine what really happened.

However, based on the available boring log, the upper soils consist of weak and likely sensitive 

clays (see appendix A). These clays are probably the source of the fill stability problems observed

at this site. In any case, the bridge abutment at site 10 appeared to be unaffected by the observed 

approach fill movement, although the superstructure for one of the bridges at this site collapsed.

Before 

After 

Shear wall 

failure

Severe fill 

settlement
99

Source: MOP

Figure 190. Photo. View of roadway fill failure from abutment 

of Route 5 railway crossing at Hospital.

Source: MOP

Figure 191. Photo. Close-up of approach fill failure and damaged sign bridge.

5.2.2 Liquefaction Impact on Performance of Bridge and Pier Structures

All of the following general conditions are necessary for liquefaction to occur:

The presence of groundwater, resulting in a saturated or nearly saturated soil.

Predominantly cohesionless soil that has the right gradation and composition.

Liquefaction has occurred in soils ranging from low plasticity silts to gravels.

Clean or silty sands and non-plastic silts are most susceptible to liquefaction.
100

A sustained ground motion that is large enough and acts over a long enough period of 

time to develop excess pore-water pressure equal to the effective overburden stress, 

thereby significantly reducing effective stress and soil strength.

Soil characterized by a density that is low enough for the soil to exhibit contractive 

behavior when sheared undrained under the initial effective overburden stress.(13,14)

Since detailed, site-specific subsurface data were not available for most of the sites visited, heavy 

reliance was placed upon visual surficial evidence that liquefaction had occurred to help interpret 

the structure performance observed. This section summarizes examples of those observations, how

the liquefaction issue has been dealt with in the past, and the observed performance of various 

structures. That performance is interpreted with consideration of the local geology (and subsurface

and ground water data if available), the likely extent of liquefaction, and the effect of the local 

topography on soil stability and deformation.

5.2.2.1 Signs of Liquefaction Observed

As previously mentioned, 15 of the bridge sites visited exhibited signs of liquefaction resulting 

from the offshore Maule earthquake. These liquefaction signs included lateral ground movement 

and cracking (i.e., lateral spreading or flow failure), ground settlement, and sand boils or otherwise

ejected sand or silt in fissures. Structures at such sites tended to display the most damage and 

the poorest seismic performance due to the heavy demands such conditions placed on them, 

with a few notable exceptions. In some cases, surface expressions of liquefaction could not be 

directly observed (e.g., piers located in rivers), so liquefaction had to be inferred based on bridge 

performance observations such as settlement or tilted piers and columns that could be explained 

in no other way.

Examples of liquefaction signs and effects are provided in figure 192 through figure 199. Figure 192

shows lateral spreading that occurred at the Mataquito bridge site (site 15). Lateral movements of 

up to 7 to 8.2 ft (2 to 2.5 m) and vertical settlement of up to 3 ft (1 m) (average of 1.6 ft (0.5 m)

across embankment near bridge abutment) appear to have occurred at the northeast end of the

structure, involving the abutment fill and over 300 ft (100 m) of gently sloping ground to the river

bank. However, as can be seen in figure 193, the lateral deformation directly beneath the bridge 

was less than the deformation just outside the edges of the bridge, indicating that the bridge 

foundations may have helped to restrain lateral movement of the soil. This also means that lateral 

soil loads may have been acting on the bridge foundations. Similar lateral spreading failures were

observed at the opposite end of this bridge, although in that case, the abutment fill was not involved

in the lateral spreading. Based on surface observations, it appears that these lateral spreading failures

at the Mataquito bridge are relatively shallow in depth, probably on the order of 7 to 8.2 ft (2 to 

2.5 m). A similar example of relatively shallow lateral spreading failure is provided in figure 194

for Ramadillas bridge (site 24), where the buildup of soil behind the bridge pier can be observed.

Based on the pattern in the ground cracks, it appears that the soil was fluid enough to attempt to 

flow around the pier but was only partially successful in doing so.
101

1 m = 3.28 ft 

Figure 192. Photo. Lateral spreading at Mataquito bridge.

Figure 193. Photo. Lateral spreading at Mataquito bridge with lateral movement 

beneath and around bridge.

Direction of 

movement

Approx. 1.5 to 1.8 m 

high

Approx. 2 to 2.5 m 

lateral movement

Lateral spreading lobes 

“pooched out” on both 

sides of bridge, but not 

directly under bridge
102

Figure 194. Photo. Lateral spreading and ground failure at old Ramadillas bridge.

Figure 195 and figure 196 provide examples of sand boils and ejected sand from ground fissures 

due to liquefaction. In figure 195, the southwest end of the Llacolen bridge, the ground water depth

is relatively shallow (less than 3 to 7 ft (1 to 2 m)). In figure 196, La Mochita bridge, the ejection of 

sand through ground fissures due to elevated pore pressures from liquefaction appeared to be more

vigorous, considering the greater depth to the water table (approximately 13 to 16 ft (4 to 5 m)). 

Figure 195. Photo. Sand boils at southwest end of Llacolen bridge.

Lateral spreading 

flow directions

Sand boils
103

Figure 196. Photo. Ejected sand due to liquefaction and ground failure at La Mochita bridge.

Figure 197 and figure 198 provide examples of more severe ground failure caused by liquefaction.

In these cases, the ground water needed for liquefaction was 7 to 16 ft (2 to 5 m) below the 

ground surface, and the remaining soil above that layer was non-liquefied crust. Due to the 

weakness of the liquefied soil, a slope failure ensued. This type of failure can cause the greatest 

damage to bridges due to the significant volume of stronger, heavier soil moving laterally into 

the bridge foundations and vertically along the foundations. Because the non-liquefied crust is 

stronger than the liquefied soil, it is less able to flow around the foundations and is more likely to 

force the foundations to bear the full brunt of the lateral soil force.

Ejected sand due 

to liquefaction

Ground 

failure

Direction of 

movement
104

Figure 197. Photo. Severe ground failure at La Mochita bridge.

Figure 198. Photo. Lateral spreading and ground failure at Raqui 2 bridge.

Figure 199 shows an example in which a center pier in the middle of the river settled (about 1.9 to

3.1 inches (50 to 80 mm)), but, because of the water in the river, no specific evidence of liquefaction 

could be observed. It is likely that liquefaction is the cause of this kind of effect on the bridge, 

although it is also likely that the liquefied layer is not extensive because the amount of deformation 

in the foundation is relatively small.

Graben

Reverse 

scarps

Direction of ground 

movement

Pier settled and 

moved laterally

Slope crest

Fill settlement relative 

to bridge pier

Lateral spreading of 

approach fill
105

1 cm = 0.39 inches

Figure 199. Photo. San Nicolás bridge center pier settlement probably 

due to liquefaction of soil layers beneath river.

5.2.2.2 Use and Performance of Ground Improvement to Mitigate Liquefaction Effects

According to engineers at MOP and Chilean universities, bridges and other transportation 

structures have historically not been designed to resist the effects of liquefaction. Until the 2010 

offshore Maule earthquake, liquefaction had only been reported in limited areas of hydraulically 

placed fill, and the problems resulting from that liquefaction had been minimal. This was not the 

case for the Maule earthquake. Liquefaction problems were widespread and, in some cases, resulted

in significant failures of structures and other facilities (see GEER for additional information).(2)

The only ground improvement the team was made aware of to address liquefaction was deep 

dynamic compaction (DDC) conducted beneath some of the approach fills near the north end of the

Juan Pablo II bridge in Concepción. Based on field observations at that site, the ground improvement

was only partially successful, as shown in figure 200 and figure 201. DDC is generally not very 

effective for stabilizing deeper liquefaction in soils below the water table (i.e., water is virtually 

incompressible, and, in finer sandy or silty soils, the water does not have time to escape during 

DDC impacts, preventing reduction of pore volume). However, the specific limits of the DDC

and the subsurface conditions are not known, so the apparent lack of success of the DDC cannot 

be fully investigated at this time.

Pier settlement of 

5 to 8 cm
106

Figure 200. Photo. Ramps at Juan Pablo II bridge where existing ground 

was improved with DDC.

Figure 201. Photo. Lateral movement of fill in ramps at Juan Pablo II bridge.

5.2.2.3 Site-Specific Impacts of Liquefaction on Bridge Performance

Among the sites visited by TIRT, the impact of liquefaction on bridges and their approaches 

varied from no impact to structure collapse. The specific reasons why some bridges performed 

reasonably well despite signs of liquefaction cannot be determined without subsurface data to 

assess the depth of liquefaction and what should have been expected regarding settlement and 

lateral spreading or flow failure. However, it is possible to summarize the damage trends that are 

likely the result of liquefaction effects.

Liquefaction can result in vertical and lateral bridge foundation movements that exceed a bridge’s

ability to tolerate them as well as stresses in foundation and substructure elements that can result 

Vertical 

settlement 

of fill

Ground cracks 

and settlement

Lateral movement of 

fill has compressed and 

crumpled guardrail

Ground 

cracking in slope
107

in component damage or failure. Liquefaction can reduce lateral soil stiffness (i.e., p-y curves),

create downdrag loads on foundation elements due to liquefaction settlement, cause permanent

lateral displacement of foundation and substructure elements, and increase lateral forces acting 

on foundation and substructure elements, which can cause the element to crack or fail. If ground

improvement to prevent liquefaction is not conducted, the foundation elements must be designed 

to resist the loads placed on them due to liquefaction.

Specific descriptions of the effect liquefaction may have had on bridge performance for several 

of the sites visited are provided in the sections that follow.

5.2.2.3.1 Mataquito Bridge (Site 15): The Mataquito bridge provides an example of the 

performance impact caused by a relatively shallow lateral-spreading failure. In this case, the depth

of lateral spreading appeared to be approximately 4.9 to 8.2 ft (1.5 to 2.5 m) based on the surficial

expression of the soil movement. The bridge substructure/foundations were not significantly 

pushed laterally, and the affected bridge spans remained intact and functional. The Ramadillas

bridge, though not discussed in detail here, provides a similar example of the effects of shallow 

lateral spreading (see figure 194).

For the Mataquito bridge, it appears that the abutment fill primarily moved parallel to and around 

the abutment wall rather than applying the full lateral force of the soil directly into the abutment. 

Ridging along the approach fill slope toe and longitudinal roadway pavement cracking appear to 

confirm that most of the soil movement was perpendicular to the roadway centerline (also identified 

in GEER).(2) Significant abutment fill settlement (approximately 1.6 to 3 ft (0.5 to 1 m)) also

occurred, requiring some easily accomplished repair to the bridge approach fill (see figure 202 and

figure 203). It also appeared that the soil was able to partially flow around the pier and foundation 

elements, possibly reducing the lateral loads enough that the foundations could resist them (see 

figure 192 and figure 193). As shown in figure 202 and figure 203, the bridge itself was virtually 

unaffected by the lateral spreading other than some minor abutment wall crushing where the

abutment and bridge deck were shoved together either during shaking or as a result of very small 

lateral movement of the abutment. There was not enough movement or damage to affect bridge 

operation. Settlement of the soil around the interior piers was also observed (see figure 204) and 

appeared to be partly due to lateral spreading of the soil down slope as well as to possible deeper 

liquefaction-induced settlement that contributed only to downdrag forces on the foundations. 
108

1 m = 3.28 ft

Figure 202. Photo. Abutment fill settlement of northeast abutment of Mataquito bridge.

1 m = 3.28 ft

Figure 203. Photo. Soil settlement without bridge settlement at northeast abutment 

of Mataquito bridge.

Fill settlement 

(approx. 0.7 to 

0.9 m)

Direction of soil movement

Soil appears to be 

pushed up here 

(0.2 to 0.3 m)

Abutment fill 

settlement (0.5 m 

average, 1 m 

max.)

Minor 

abutment wall 

compression 

damage
109

1 m = 3.28 ft

Figure 204. Photo. Soil settlement at interior pier at northeast abutment 

of Mataquito bridge.

Specific details regarding the foundations of the bridges were not available, but considering the 

recent construction date for the bridge (2007–2008) and the size of its columns, it is likely that the

bridge was founded on drilled shafts (typical diameter of 4.9 ft (1.5 m)) that are deep enough to 

reach denser soil. This may explain why the bridge itself did not settle, but that hypothesis will 

need to be confirmed once bridge foundation details and soil subsurface information are available.

5.2.2.3.2 Raqui 1 bridge (Site 20): The Raqui 1 bridge is somewhat transitional with regard to 

the effects of liquefaction. The primary impact of liquefaction for this bridge was abutment fill 

settlement, though evidence of relatively minor lateral spreading below the abutment fills was

present. Figure 205 shows approximately 3 ft (1 m) of fill settlement at the northwest abutment.

The southeast abutment shows a similar amount of settlement. However, figure 206 shows that some

lateral movement of the northwest abutment may have occurred (approximately 2.9 to 3.9 inches

(75 to 100 mm)), severe enough to cause crushing of the abutment wall against the bridge girders 

and deck but not to shut down the bridge once the abutment approach fills were rebuilt to match 

the bridge deck elevation. A small concrete gravity wall adjacent to the abutment also displaced 

and tilted laterally down slope toward the river (see figure 206). It also appeared that more lateral 

movement occurred on one edge of the northwest abutment than on the other, indicating that some

twisting of the abutment occurred. The abutment wall crushing could be partially due to strong

shaking of the superstructure, forcing the bridge girders and deck into the abutment. However, the

abutment wall crushing, signs of abutment fill displacement, twisting of the abutment, and lateral 

movement of the gravity wall adjacent to the abutment indicate that abutment movement toward 

the river occurred, likely due to liquefaction. It appeared that the southeast abutment did not move

due to the liquefaction and lateral spreading. The difference in the liquefaction effects at the two 

abutments may be due to the presence of more gently sloping ground around the approach fill for 

the southeast abutment than for the northwest abutment (see figure 207). There was some evidence

0.5 m
110

that the direction of slope movement was perpendicular to the roadway centerline, especially a few

meters or more away from the abutment (see figure 208). However, since the approach fills had 

been restored to the bridge deck elevation before the team’s arrival, some of the slope failure 

cracking and movement patterns had been covered by new fill. The age of the bridge, the 

subsurface conditions, and the foundation details were not available to TIRT. Therefore, 

additional investigation is needed to more accurately assess the bridge performance.

Figure 205. Photo. Abutment fill settlement at northwest abutment of Raqui 1 bridge.

Fill settlement

Girders jammed 

against abutment 

wall
111

Figure 206. Photo. Abutment and girders shoved into each other and deck at the northwest 

abutment of the Raqui 1 bridge. 

Figure 207. Photo. Raqui 1 bridge viewed from north side.

Girders/deck 

jammed against 

abutment wall

Concrete gravity wall 

tilted and displaced
112

Figure 208. Photo. Evidence of ground failure perpendicular to roadway centerline 

at northwest approach of Raqui 1 bridge.

5.2.2.3.3 Llacolen Bridge (Site 17): Figure 195 illustrates another case where liquefaction 

obviously occurred (sand boils and evidence of minor lateral spreading). In this case, the 

southwest end of the Llacolen bridge, the liquefaction had very little impact on the bridge both

in terms of foundation settlement and lateral movement of the foundations and abutment.

At the northeast end of the Llacolen bridge, the liquefaction impacts were far worse. The biggest 

problem for the Llacolen bridge, as well as for other Biobío River crossings and structures, was 

lateral movement of the foundations within the liquefaction-induced slope failure areas (see 

figure 209 for an overview of the river crossings and figure 210 through figure 213 for the 

Llacolen bridge). In general, most of the liquefaction-induced slope failure problems occurred 

along the northeast side of the river, where the river channel is incised more steeply. Very few 

liquefaction-induced problems occurred along the southwest side of the river even though there 

was clear evidence that liquefaction had occurred, probably due to the very flat ground 

conditions on that side of the river (see figure 195).

Lateral ridging due 

to ground failure

Direction of 

movement
113

© Google, Inav/Geosistemas SRL, DMapas, and GeoEye

Figure 209. Map. Aerial view of Biobío River and crossings.

Figure 210. Photo. Evidence of slope instability due to liquefaction 

at northeast abutment and approaches of Llacolen bridge.

Direction of river 

flow Observed liquefaction induced 

slope failures

Railroad 

crossing

La Mochita 

Bridge

Llacolen

Biobío

Juan 

Pablo II

Ground cracks 

and settlement 

observed in 

these areas
114

1 m = 3.28 ft

Figure 211. Illustration. Plan view of liquefaction-induced slide at northeast abutment

of Llacolen bridge.

Figure 212. Photo. Settlement due to liquefaction-induced slope instability

at northeast end of Llacolen bridge.

25 m

Puente

Llacolen

Direction of movement 

(lateral deformation

approx. 0.25 to 0.3 m)

Settlement of 

0.4 m

Collapsed

span

Cracked columns,

and foundation 

pushed outward

30 m

9 m 4 m

40 m

20 m

River’s edge

Walkway

50 m

Roadway

Head-scarp

0.4 m high

Ramp

Approximate 6 to 8 m 

elevation drop 

from walkway

to river

Toward

south

abutment

Ground cracks 

and settlement

Gaps present 

in front of pier 

shafts
115

1 m = 3.28 ft

Figure 213. Photo. Settlement and lateral ground movement of interior pier

at northeast end of Llacolen bridge.

At the northeast end of the Llacolen bridge, the liquefiable layer was probably near the river 

level. As the liquefiable layer weakened during shaking, the weight of the non-liquefiable soil 

caused the strength of the liquefied soil to be exceeded, resulting in a slope failure. As the soil 

moved laterally and downhill, it increased lateral stresses acting on the foundations, pushing 

them downhill and increasing bending stresses in the columns. The lateral movement caused a 

span to fall off its seat and collapse as well as excessive bending stresses to be placed on the pier 

columns, causing the columns to crack at a point of maximum stress in combination with 

weakness caused by a construction joint at that location. Similar observations and conclusions 

were made by the GEER team.(2) This happened for several of the columns shown in figure 214

and figure 215. It is not completely clear whether the columns cracked as a result of lateral 

movement of the slope caused by liquefaction or if once the pier moved and the span fell, the 

impact of the falling span on the ground immediately behind the cracked columns was the final 

effect that cracked the columns. In either case, the lateral ground movement significantly 

increased the lateral stress on the bridge foundations and columns.

Soil settlement of 0.4 m 

and lateral movement 

of 0.25 m

Toward river

Fallen span
116

Figure 214. Photo. Lateral ground and foundation movement at interior pier and fallen 

span at northeast end of Llacolen bridge.

Figure 215. Photo. Close-up of cracked column at northeast end of Llacolen bridge.

The shaft foundations supporting the bridge were approximately 98 ft (30 m) deep and were 

probably bearing in dense soil. This kept the liquefaction settlement from dragging the foundations 

down, limiting the liquefaction impact to lateral movement of the foundations. This was not the 

case for the Juan Pablo II bridge, where shallower (59-ft (18-m)-deep) caisson foundations that 

Slope cracks
117

probably did not bear on dense soil were used. Foundation settlement did occur for the Juan 

Pablo II bridge, though only at specific piers (see discussion in section 5.2.2.3.5).

Additional photos and details regarding the Llacolen bridge and the impacts of liquefaction on 

that bridge are provided in GEER.(2)

5.2.2.3.4 Chepe Railroad Bridge Over Biobío River (Site 18): A railroad bridge built in 1889 near

and parallel to the Llacolen bridge also suffered damage to one of its piers due to lateral movement

and settlement caused by liquefaction in the vicinity of the northeast river bank. It appeared that 

the 7-ft (2-m)-diameter steel pipe pile concrete-filled foundation and column displaced laterally 

about 2 ft (0.7 m) and that the ground around the foundation settled approximately 4.3 ft (1.3 m) 

(see figure 177). The roadway that this bridge crosses also settled approximately 4.3 ft (1.3 m) 

and moved laterally approximately 1.6 to 2 ft (0.5 to 0.7 m), though the wall top also rotated

inward towards the roadway (see figure 216). The wall itself appeared to settle only about 1 to 

1.3 ft (0.3 to 0.4 m). The roadway and walls formed a seal to keep groundwater out of the roadway

(see chapter 4 for additional details). The liquefaction-caused settlement broke that seal, which 

will need to be repaired before the river level rises. Even with so much deformation, these 

structures did not collapse despite being seriously damaged.

1 m = 3.28 ft

Figure 216. Photo. Wall settlement due to slope movement and instability caused by 

liquefaction at northwest end of Chepe railroad bridge.

5.2.2.3.5 Juan Pablo II Bridge (Site 25): Juan Pablo II bridge, another crossing of the Biobío River 

parallel to Llacolen bridge in Concepción, was built in 1973. This bridge was also impacted by 

liquefaction-induced slope failure as well as liquefaction-induced settlement and downdrag.

The 1973 bridge foundations consisted of 59-ft (18-m)-deep hand-excavated caissons. Severe 

liquefaction-induced settlement of foundations occurred at several locations across the length of 

the bridge (1.6 ft (0.5 m) or more), but many of the piers exhibited no discernable settlement.

Roadway settled 

approx. 1.3 m

Railing/wall 

displacement
118

Notably, a shaft-supported wall supporting fill between the approach bridge and the main rivercrossing structure settled very little, whereas the bridge approach pier settled more than 3 ft (1 m),

and the soil surrounding the end of the wall and the bridge pier also settled more than 3 ft (1 m)

(see figure 217 and figure 218). This difference is likely due to differences in the two structures 

and the amount of load they carry. It was noted that the shaft-supported wall is much newer than 

the bridge ramp structure and therefore may be designed to more stringent standards, as it is 

unlikely that the soil would be so vastly different over such a short distance. Figure 219 through

figure 221 show the bridge settlement.

Figure 217. Photo. Settlement of approach structure for Juan Pablo II bridge.

Figure 218. Photo. Close-up of settlement of approach structure for Juan Pablo II bridge.

South wall

Pier supporting 

approach structure 

settled, but wall did not

Bridge settlement

South wall
119

Figure 219. Photo. Roadway settlement due to settlement of approach structure 

pier foundations for Juan Pablo II bridge.

Figure 220. Photo. Roadway settlement due to settlement of mainline pier foundations 

for Juan Pablo II Bridge.
120

Figure 221. Photo. Liquefaction-induced differential settlement of mainline 

pier foundations supporting Juan Pablo II bridge.

The mainline Juan Pablo II Bridge abutment at the river bank suffered damage due to the lateral 

forces caused by liquefaction-induced slope instability. Figure 222 and figure 68 through figure 71

illustrate the damage that occurred. The damage to the column in figure 68 through figure 71

illustrates the severe bending stresses and lateral force to which the column was subjected when the 

northeast riverbank slope moved laterally due to loss of soil strength caused by liquefaction.

Although the bridge did not collapse due to these severe differential settlements and lateral slope 

movements, it was seriously damaged and efforts were already underway to design and build a 

replacement structure. Preparations for test hole drilling were being made during the TIRT visit.

Figure 222. Photo. Damage to pier columns for mainline Juan Pablo II bridge.

Failed column

Direction of 

slope 

movement
121

Note that GEER also provided a detailed account of reconnaissance of this structure and the 

damage that occurred due to liquefaction impacts. Additional photos and measurements for that 

structure are provided in GEER’s report.(2)

5.2.2.3.6 Old Biobío River Bridge (Site 26): The old Biobío River bridge was built in the 1920s,

replaced in 1943, and no longer in service at the time of the earthquake. Spans and some of its 

substructure units collapsed in the earthquake, probably due to a combination of inadequate structural

resistance and shallow foundations (see figure 138, figure 223, and figure 224). Various types of 

deep foundations were used in various sections of the bridge. It appeared that pile foundations were

used for the northeast half of the bridge, although it was not clear if all the original timber pile 

foundations used to support the 1920s structure were replaced with steel pipe piles in the 1943 

replacement structure. Therefore, the northeast portion of the bridge could be supported by both 

timber and steel pipe piles. Based on input from a Chilean consultant, the piers near the southwest 

end of the bridge appeared to be supported by twin shaft foundations or twin hand-dug oval caissons 

(33 to 39 ft (10 to 12 m) deep, although it is possible that the foundations were tipped in relatively

dense soils). Scour problems and deep liquefaction or soft soils that allowed more foundation 

movement during shaking may have contributed to the weakening and collapse of the structure, 

though the specific role of liquefaction in this collapse cannot be confirmed until more subsurface

data becomes available. However, ground failure and lateral soil movement were observed at the 

bridge abutment for this structure on the northeast bank of the Biobío River, and the abutment 

pier wall appeared to be tilted toward the river (see figure 137, figure 225, and figure 226).

Figure 223. Photo. View of collapsed spans from northeast river bank of Biobío River.

Figure 224. Photo. Close-up of old Biobío River bridge showing exposed pile foundations.

Exposed piles
122

Figure 225. Photo. Ground failure near northeast abutment of old Biobío River bridge.

Figure 226. Photo. Tilting of northeast abutment pier wall of old Biobío River bridge.

5.2.2.3.7 La Mochita Bridge (Site 27): Significant ground failure due to liquefaction also occurred

at the La Mochita bridge, located beside the Biobío River just south of the three Biobío River 

crossings. The north approach fill to this structure and the soil surrounding the interior piers 

exhibited the most severe ground failure. Figure 227 through figure 229 show examples of the 

Outward rotation 

of abutment wall

Lateral movement of 

girders
123

north approach fill damage, and figure 196 and figure 197 show ground failure at the interior piers. 

The presence of ground water is a key ingredient for liquefaction to occur, and the ground water 

regime at this site was complicated by the presence of a dredging operation that pumped water 

from the river into a pit adjacent to the north approach fill, possibly raising the ground water levels

in that approach fill (see figure 229). Seepage, possibly from water in the dredging spoils pit, was

observed in the riverbank slope about 7 to 10 ft (2 to 3 m) above the river level in this area. This may

have increased the severity of the ground failure observed in the north approach fill. However, 

the severity of the ground failure did not seem to be as great at the bridge abutment as it was 

82–660 ft (25–200 m) north of the abutment (see figure 227 and figure 228). This may have 

been one of the reasons why the north bridge abutment was not significantly displaced due to 

this liquefaction-induced soil movement.

Figure 227. Photo. Ground failure of north approach fill for La Mochita bridge.
124

Figure 228. Photo. La Mochita bridge north approach fill liquefaction-induced 

ground failure.

Figure 229. Photo. La Mochita bridge north approach fill and edge of dredging pit.

Ground failure due to liquefaction at the bridge location was most obvious at the interior piers (see

figure 196 and figure 197). While it appeared that the direction of movement was generally 

Puente La Mochita

Dredging/sluicing 

operation

Biobío River
125

eastward, away from the main river channel, the presence of reverse scarps and vertical movement

of blocks of soil appeared to indicate the formation of a graben at and just west of the piers. This

indicates that most of the ground movement at the interior piers was vertical with limited lateral 

movement. This may explain the relatively small lateral movement of the interior piers in spite of 

the apparently larger lateral slope movements. In any case, as can be seen in figure 197, the interior 

pier settled and tilted eastward, away from the river. The interior piers are believed to be supported 

on four 4.9-ft (1.5-m)-diameter drilled shafts, but this needs to be verified. Figure 230 shows the 

effect of the pier settlement on the elevation of the bridge deck.

Figure 230. Photo. La Mochita bridge deck and girder lateral movement.

The bridge superstructure (girders and deck) also slid laterally toward the east up to 1 ft (0.3 m), 

with more movement at the south abutment than at the north abutment, possibly due to a combination

of strong shaking and the tilting of the interior pier (see figure 230 through figure 232). It appeared

that the abutments did not move. Note that one of the outside bridge girders almost slipped off the

abutment seat (see figure 232). Overall, while the ground failure was significant (lots of movement

with a thick non-liquefied crust), this bridge did not collapse. However, it was laterally displaced 

and suffered damage.

Minor bridge deck 

settlement

Lateral bridge deck 

movement
126

Figure 231. Photo. La Mochita bridge deck and girder lateral movement 

and approach fill settlement.

Figure 232. Photo. La Mochita bridge lateral bridge deck and girder at south abutment.

Additional observations and measurements regarding the ground failure and bridge movements, 

including photos, are provided in GEER. The overall conclusions made by GEER are consistent 

with the observations of TIRT.(2)

Lateral bridge deck 

movement

Girder almost slipped 

off bridge deck

Lateral movement of 

girders
127

5.2.2.3.8 Raqui 2 and Tubul Bridges (Sites 21 and 22): Examples of severe bridge damage due to

liquefaction combined with inadequate structural resistance include the Raqui 2 and Tubul bridges.

For the Raqui 2 bridge, damage caused by liquefaction included severe failure and large-scale 

settlement of the approach fills (see figure 198 and figure 233) and collapse of the interior bridge 

spans resulting from movement and tilting of the interior pier foundations (see figure 234 and

figure 235). It was not clear if the lateral displacement of the interior piers was due to some type 

of lateral soil movement or permanent displacement due to the combination of weak liquefied soils 

and lateral inertial forces induced by shaking of the bridge superstructure. Even though severe 

failure of the abutment fills occurred (likely due to liquefaction), it appeared that the abutments 

themselves did not move significantly. Based on the lateral ridging observed along the approach 

fill side slope toes and the roadway cracking pattern, it appeared that most of the lateral soil

movement was in the direction perpendicular to the roadway centerline, which could explain why

the abutments did not appear to have moved laterally.

Figure 233. Photo. Southeast approach fill damage at Raqui 2 bridge.

Approximately 4.9 ft 

(1.5 m) of approach fill 

settlement
128

Figure 234. Photo. Collapsed and shifted spans at Raqui 2 bridge as viewed 

from base of southeast abutment.

Figure 235. Photo. Collapsed and shifted spans at Raqui 2 bridge due to liquefaction.

Damage caused by liquefaction for the Tubul bridge was similar to that observed at the Raqui 2

bridge, though more lateral movement of the abutments due to lateral spreading forces appeared to

have occurred. Ground cracking and deformation indicative of lateral spreading toward the river 

Span has shifted to left 

(SW) while pier wall 

did not move

Tilted substructure pier 

(may have tilted during 

shaking due to liquefaction 

or when span fell)
129

was observed at both ends of the bridge. Damage observed included liquefaction-induced ground 

failure, large-scale settlement of the approach fills (see figure 236 through figure 240), and shoving 

of the bridge girders approximately 5.9 inches (150 mm) into the north abutment, resulting in some 

crushing of the abutment wall at the girder locations (see figure 240 and figure 115). The south 

abutment pile cap was also damaged (see figure 239). It is possible that the north abutment was 

pushed into the girder to cause the abutment wall crushing. Alternatively, the girders supported by 

the interior piers could have been shoved into the north abutment due to a combination of lateral 

movement of the south abutment and strong shaking (see chapter 4).

Figure 236. Photo. South approach fill damage and collapsed spans at Tubul Bridge.

Figure 237. Photo. Settlement of fill beside the south abutment of Tubul bridge.

Approach fill 

settlement
130

Figure 238. Photo. Overview of damage to south abutment of Tubul bridge.

Figure 239. Photo. Close-up of exposed pile and pile cap damage 

on south abutment of Tubul bridge.

Timber pile 

and damage to 

pile cap

Timber pile 

and damage to 

pile cap
131

Figure 240. Photo. Overview of Tubul bridge north abutment.

The exposed pile cap of the Tubul bridge south abutment was obviously damaged, possibly due to

lateral movement toward the north (see figure 239). It is not clear if the exposed timber pile was 

embedded in the pile cap as part of the abutment foundation or was a remnant of a foundation 

from an older bridge that ended up getting included in the pile cap when it was constructed.

Specific foundation details for this bridge were not available, although it is likely that this 

bridge was supported on piles. In any case, it appears that the presence of the timber pile in 

combination with lateral movement of the abutment due to lateral spreading forces may have 

forced the exterior concrete pile cap to break and come loose. Additionally, the approach fill 

settlement and lateral movement appeared to be more severe at the south abutment than at the 

north abutment, providing further support that the collapse may have initiated at the south end of 

the bridge. Regarding the interior piers of the Tubul bridge, damage or collapse of the pier walls 

appeared to be the primary problem. The pile caps appeared to not have moved, at least in the 

lateral direction. However, some of the pile caps were not visible (they were under water or mud)

and therefore could have settled or moved laterally (see figure 241).

Figure 241. Photo. Tubul bridge collapsed spans and substructure.

Approach fill 

settlement/movement

Abutment and girder 

were shoved together 

Cracks due to lateral 

spreading
132

The age of these bridges, the subsurface conditions, and the foundation details were not available

to TIRT. Therefore, additional investigation is needed to more accurately assess the observed 

bridge performance.

5.2.2.3.9.Puerto de Coronel Muelle (Site 19): Deep liquefaction combined with lateral soil 

movement can cause severe foundation element bending. This was exemplified at the Puerto de 

Coronel Muelle, located south of Concepción but north of Arauco. The foundations supporting 

the pier (wharf) structure were pushed laterally and displaced (see figure 242 and figure 243).

Since the foundations were deep (steel 28- to 40-inch (711 to 1,016 mm) open-ended pipe piles 

driven to a depth of approximately 160 ft (50 m)), the piles were probably fairly well fixed near 

the bottom. Based on the boring logs available, loose sands are present in the top 49 to 66 ft (15 

to 20 m) below the ground surface. Therefore, liquefaction likely occurred to significant depth 

and contributed to the foundation displacements and degrees of bending observed. The yellow 

line in figure 243 traces the bending and twisting observed in the girder that was attached to the 

pile foundation. See GEER for additional details regarding the effect lateral spreading had on 

these and similar structures in the vicinity.(2)

Figure 242. Photo. Puerto de Coronel Muelle foundation displacement and bending.

Figure 243. Photo. Close-up of Puerto de Coronel Muelle foundation displacement and bending.

Lateral soil movement 

forced pile over
133

5.2.2.4 Summary of Liquefaction Performance for Bridges and Other Structures: In summary,

the primary effects of soil liquefaction on bridges and pier structures included foundation and 

approach fill settlement, loss of lateral foundation support, and lateral foundation movement due 

to lateral spreading and liquefaction-induced slope failure. Foundation settlement in particular 

was highly variable, ranging from 0 to 3 ft (0 to 1 m) or more, even for piers that were next to 

each other (e.g., Juan Pablo II bridge). Foundation settlement depends not only on the thickness 

of non-liquefiable crust (affecting downdrag unit friction on the foundations) and thickness of 

liquefiable soil (affecting magnitude of soil settlement) but also on the ability of the soil at the 

foundation tips to resist deformation and provide the needed bearing resistance. Note that even 

with such severe differential settlement, bridge collapse did not occur.

Lateral displacement of abutment walls and interior piers appeared to occur primarily when a 

significant crust of non-liquefied soil slid on top of a liquefied soil layer and when the slope was 

wide enough to significantly reduce or eliminate three-dimensional effects. Table 5 summarizes 

the approximate slope geometry and performance of bridge abutments and interior piers. The table

entries are in general order of increasing severity of soil and structure movements and damage.

The information in the table demonstrates that abutments that form the terminus of a long, narrow

fill tend to be mostly unaffected by liquefaction and lateral ground movement, assuming that the 

abutment is supported by a reasonably robust foundation. In general, site observations confirmed 

that most of the soil movement was perpendicular to the roadway fill centerline and not directly into

the abutment wall. Overall, bridge abutments did well in this earthquake, even when liquefaction 

and resulting ground settlement or failure was severe. However, where the bridge abutment, interior

pier, or pier structure foundations were located on a broader slope such as would occur adjacent 

to a river or ocean waterfront, liquefaction-induced ground failure and lateral soil movements did 

result in lateral displacement or overstress and failure of the foundations and substructure. A 

possible reason for this difference in performance is the ability of the soil to either flow around 

the foundations or to otherwise follow the path of least resistance or greatest driving force. This 

depended on the thickness and fluidity of the failed soil mass and the slope and bridge geometry.

The greatest driving force is likely to be in the direction where three-dimensional effects have 

the smallest influence and smallest contribution to the resistance to movement.
134

Table 5. Summary of performance for abutments and interior piers on slopes affected by liquefaction-induced 

ground movement or settlement.

Site No., Name, and 

Abutment/Pier Location

Estimated 

Abutment 

or Slope Height

Estimated Ground Geometry 

at Abutment/Pier Vicinity

Observed Abutment/Pier 

Deformation and Performance

17. Llacolen bridge, SW 

abutment and land side piers 1 to 2 m ~ 3-degree slope (see figure 195) No movement or significant damage

20. Raqui 1 bridge, SE 

abutment 3.5 m

Two-lane approach fill on gently sloping ground 

(see figure 207)

No movement/significant damage to abutment. 1 m of 

approach fill settlement and lateral spreading 

perpendicular to the roadway centerline.

15. Mataquito bridge, both 

abutments 7 to 8 m

Gently sloping to river (~ 5 degrees) (see figure 192, 

figure 193, and figure 202 through figure 204)

NE abutment shows minor crushing on abutment wall. If 

lateral movement occurred, it was less than 20 mm. SW 

abutment did not move

27. La Mochita ridge, both 

abutments

9 to 12 m transverse 

to abutment toward 

river

Steeply sloping (~ 25 degrees) to river transverse to 

bridge abutment wall; rip-rapped ~ 25-degree slope 

below abutment to river-level north abutment and to 3 to 

4 m high ridge below bridge centerline at south abutment 

(see figure 197 and figure 228 through figure 232).

No abutment movement or significant damage; fill 

settlement at abutments varied from 0.3 m to 0.8 m. 

Lateral movement of approach fills perpendicular to 

roadway centerline.

20. Raqui 1 bridge, NW 

abutment 4 m

Moderately sloping to river (~ 7-degree slope) 

(see figure 207)

Abutment fill settled about 1 m. Abutment may have 

moved laterally 75 to 100 mm and twisted. Small 

adjacent wall parallel to abutment wall tilted and moved 

laterally.

27. La Mochita bridge, 

interior pier 3 to 5 m

Ground slopes perpendicular to bridge centerline and 

away from main river channel (see figure 197)

Pier moved laterally about 100 to 150 mm, and vertically 

about 50 to 100 mm. Movement was down slope away 

from Biobío River main channel.

19. Puerto de Coronel Muelle, 

pier on sloping shoreline 4 to 5 m General slope toward ocean at ~ 10 degrees

Pile foundations moved laterally down slope about 0.5 to 

0.7 m.

21. Raqui 2 bridge, both 

abutments and interior piers 4 m

Two lane approach fill on gently sloping ground (see 

figure 198, figure 233, and figure 234)

No abutment movement but some minor damage, 

approach fill settlement up to 1.5 m, and lateral 

spreading perpendicular to the roadway centerline. 

Interior spans collapsed and piers tilted laterally.

18. Chepe railroad bridge over 

Biobío River, NE river bank 

pier 7 to 8 m

Pier located at top of river bank; geometry is complex 

(see figure 176 and figure 216)

Pier moved laterally 0.7 m; ground around pier settled 

1.3 m.

26. Old Biobío River bridge, 

NE abutment and interior 

piers 6 to 7 m

Abutment on river bank slope (~ 25 degrees) (see figure 

225 and figure 226)

Abutment wall tilted down slope about 150 mm. 

Direction of ground movement is toward river. Some 

interior piers collapsed.

17. Llacolen bridge, pier on 

slope below NE abutment

5 to 6 m at 20-degree 

slope, then less than 

10 degrees with 2-m 

drop to river edge

Ground slopes down to river (see figure 210 through 

figure 214)

Pier on river bank moved laterally about 0.25 to 0.3 m 

and vertically about 0.3 to 0.4 m, causing columns to 

crack all the way through and approach span to move off 

bridge seat and collapse.

See notes at end of table
135

Site No., Name, and 

Abutment/Pier Location

Estimated 

Abutment 

or Slope Height

Estimated Ground Geometry 

at Abutment/Pier Vicinity

Observed Abutment/Pier 

Deformation and Performance

22. Tubul bridge, north 

abutment and interior piers 4.5 to 5 m

Fill behind abutment extends out from edges of bridge 

(wider than needed for two lanes); slope is 25 to 30 

degrees (estimated) then flattens to 10 degrees or less in 

short section to river (see figure 240)

Abutment may have moved laterally about 150 mm, but it 

is also possible that girders were shoved into abutment due 

to a combination of pier wall failures, strong shaking, 

and lateral movement of south abutment. Approach fill 

settled over 1 m. Interior piers collapsed.

22. Tubul bridge, south 

abutment 5 m

Two-lane approach fill on ~ 10-degree sloping ground 

(see figure 236 through figure 240)

Approach fill settled up to 1.5 m with lateral spreading

primarily perpendicular to the roadway centerline, but 

turning toward river away from the abutment. Front of 

pier cap was damaged at timber pile location, possibly 

due to lateral abutment movement of 100 mm or more. 

Interior spans collapsed.

25. Juan Pablo II bridge, pier 

on slope below NE abutment 6 to 7 m

Pier is located mid-slope; slope drops steeply to river 

(~ 30 degrees) (see figure 222 and figure 68 through 

figure 71)

Pier column on river bank slope moved laterally 0.3 m or 

more and column failed, but span did not collapse. 

Interior piers suffered severe differential settlement of up 

to 1 m.

1 m = 3.28 ft

1 mm = 0.039 inches
136

The current design philosophy in North America is to stabilize bridge abutments to prevent 

liquefaction, especially for the case of bridge abutments that form the terminus of relatively long 

approach fills not located on top of a broader slope (e.g., a riverbank). This design philosophy 

may be overly conservative and costly if large approach fill settlement can be tolerated for a nocollapse scenario. The approach fills can be simply and quickly rebuilt, as was done at several 

bridges observed by TIRT. It appeared that three-dimensional effects had a significant impact on 

the stability and movement of bridge abutments and piers subjected to liquefaction. Of course, 

these conclusions about liquefaction effects are preliminary until more subsurface and structure/

slope geometry data are available and analyzed. The conclusions provided herein are based on 

surficial observations, limited information about the subsurface conditions, and limited or 

visually estimated ground slope geometry.

5.2.3 Effect of Regional Ground Uplift Observed for Bridges

As discussed in chapter 2, one result of subduction zone earthquakes is regional uplift and

subsidence. The effects of regional uplift were especially prevalent south of Concepción, in the 

vicinity and south of Arauco, where regional uplift of approximately 7 ft (2 m) was observed.

This especially affected harbors and waterways, leaving ships on sand and gravel bars. It also left

beaches, estuaries, and wetlands without the water necessary to support plant and aquatic life.

Fishing villages that had close access to navigable waterways lost access, requiring new access

ways to be built.

Examples of the impact of rapid uplift are shown in figure 244 and figure 245. These figures show

dry ground that was once wet and boats hung up on old timber pilings that were once almost fully

submerged. However, the team was not able to confirm that these boats ended up on the pilings 

as the result of the earthquake. Figure 117 shows additional evidence of the pre-earthquake water 

level. One possible effect of the rapid uplift was to create rapid drawdown conditions, increasing 

the instability of sloping ground that was once partially or fully underwater. These remnant pore 

pressures could have combined with liquefaction effects to further decrease the stability of affected

slopes, making them more likely to fail and to experience greater lateral deformations. This scenario

needs further study once more subsurface data is obtained. The effect of rapid uplift on the bridge 

itself is unknown at this point. However, if the rapid uplift created the potential for differential 

movement between piers, the rapid uplift could have created additional problems for a bridge 

already in distress due to strong ground motions and liquefaction-induced forces or loss of 

foundation support. These near-fault effects need further consideration.
137

Figure 244. Photo. View northeast from Tubul bridge south abutment showing areas 

left dry and boats left above the new mean water level.

Figure 245. Photo. Tubul bridge with boats run through by old timber pilings.

Boats left above the 

mean water level due 

to regional uplift

Direction of lateral 

spreading

Boats left hanging 

on or run through 

by old piling
138

5.3 GEOTECHNICAL OBSERVATIONS REGARDING RETAINING WALL 

PERFORMANCE

In general, walls have performed very well in past earthquakes. Examples include soil nail walls 

and mechanically stabilized earth (MSE) walls in the 1989 Loma Prieta earthquake in California, 

MSE walls in the 1994 Northridge, CA, earthquake, and MSE walls and modern reinforced concrete

gravity walls the 1995 Kobe, Japan, earthquake. (See references 15–18.) Koeski, et al. provide a 

summary of the performance of various types of walls, focusing on geosynthetic reinforced

structures, in these and more recent earthquakes.(19) The largest earthquakes have shown some

differences in performance between relatively rigid walls such as concrete gravity walls and more

flexible, MSE-type walls, with the MSE type structures tending to outperform the concrete gravity

structures, especially older concrete gravity structures. While some damage to walls has occurred

in past earthquakes, the damage has typically included minor movement (4 inches (100 mm) or 

less) or outward tipping of the wall, cracking of wall facing elements, wall corners or full-height 

joints opening up and allowing backfill to spill through, unsecured blocks at the wall top toppling 

off, and on rare occasions, collapse. Instances of wall collapse in past earthquakes were almost 

always limited to older walls that would not meet current design standards (especially unreinforced

masonry and concrete walls) and, in a few cases, newer walls that were not designed to standards 

or were already in serious distress before the earthquake due to design or construction problems.

This section presents observations regarding the performance of retaining walls. Many of the 

32 sites visited had bridge embankment approaches that were either sloped embankments or 

retaining walls. In addition, many sites used retaining walls to confine the soil under the bridge

abutment. For the most part, the retaining walls observed consisted of either CIP concrete cantilever

walls or MSE walls. The MSE walls observed were constructed of either precast panels with 

metallic reinforcement or dry cast modular blocks with geosynthetic reinforcement.

Traditionally, CIP walls have been used to confine the approach fill near a bridge abutment. MSE

construction is a newer technology that has been used around the world since the 1960s as an 

alternative to CIP walls. The advantage of MSE walls include cost, rate of construction, and 

tolerance to settlement. In Chile, MSE walls have be used since 1995 (site 28 includes the first

MSE walls built in Chile). The team observed a combination of CIP and MSE walls at many of 

the sites visited. In most cases, MSE walls were only used to retain the embankment approaches 

and not under the abutment. Under the abutment, CIP concrete walls were typically used. Table 6

lists the sites where retaining walls were observed and includes brief descriptions. Figure 246

through figure 260 show photos of each of the walls.
139

Table 6. Summary of sites with retaining walls.

Site 

No. Site Name Wall Description

Wall 

Height

(m)

1

Américo Vespucio/

Miraflores

MSE wall (steel reinforcement, precast panel 

faced) approaches with CIP abutment wall 6 to 7

2

Américo Vespucio/

Lo Echeveres

MSE wall (steel reinforcement, precast panel 

faced) approaches with CIP abutment wall 6 to 7

3 14 de la Fama

MSE wall (steel reinforcement, precast panel 

faced) approaches with CIP abutment wall 6 to 7

5

Quilicura railway 

overcrossing CIP concrete abutment and approach walls 5 to 6

6a

Américo Vespucio/

Independencia eastbound

MSE wall approaches with CIP abutment walls, 

both precast panel with steel reinforcement and 

modular block with geogrid reinforcement 6 to 7

6b

Américo Vespucio/

Independencia westbound

CIP concrete back-to-back semi-gravity walls for 

approach ramps 6 to 7

9 Maipú River bridge Concrete block gravity wall 9 to 11+

11

Estribo Francisco 

Mostazal (Avenida 

Independencia)

MSE (modular block and geogrid reinforcement) 

abutment walls. Walls directly support abutment 

loads from the bridge.

6.5 to 

7.5

18

Chepe railroad bridge 

over Biobío River

CIP concrete wall supporting cut for depressed 

roadway 5

21 Raqui 2

Approach ramps use CIP walls for the embankment 

and abutment 4

25 Juan Pablo II Shaft-supported CIP walls retaining embankment fills 2 to 3

28 21 de Mayo

MSE wall (steel reinforcement, precast panel 

faced) approaches and CIP abutment walls 8 to 10

29 Rotonda General Bonilla

MSE wall (modular block face with geogrid

reinforcement) approaches and CIP abutment wall 4.5 to 5

32 Muros Talca (SW)

MSE wall (modular block facing and geogrid 

reinforcement) approaches and CIP abutment wall 8 to 9.5

1 m = 3.28 ft

Figure 246. Photo. Retaining wall at Américo Vespucio/Miraflores.
140

Figure 247. Photo. Retaining wall at Américo Vespucio/Lo Echevers.

Figure 248. Photo. Retaining wall at 14 de la Fama.

Figure 249. Photo. Retaining wall at Quilicura railroad overcrossing.

Figure 250. Photo. Retaining wall at Américo Vespucio/Independencia eastbound.

Figure 251. Photo. Retaining wall at Américo Vespucio/Independencia westbound.
141

Figure 252. Photo. Retaining wall at Maipú River bridge.

Figure 253. Photo. Retaining wall at Estribo Francisco Mostazal (Avenida Independencia).

Figure 254. Photo. Retaining wall at Chepe railroad bridge over Biobío River.

Figure 255. Photo. Retaining wall at Raqui 2.

Figure 256. Photo. South retaining wall at Juan Pablo II.
142

Figure 257. Photo. North retaining wall at Juan Pablo II.

Figure 258. Photo. Retaining wall at 21 de Mayo.

Figure 259. Photo. Retaining wall at Rotonda General Bonilla.

Figure 260. Photo. Retaining wall at Muros Talca (SW).

In general, all of the observed retaining walls performed well. There was little to no wall movement

or damage, with the exception of sites 6, 18, 25, 28, and 32. Other than these sites, the wall facing,

top-of-wall details, pavement, and other ancillary components showed no signs of distress due to 

the earthquake, even though some of the bridges associated with these walls experienced significant

damage or even complete collapse of the superstructure. Although the retaining walls at sites 6, 

18, 25, 28, and 32 experienced deformations or damage, they continued to perform their function of

retaining soil and maintaining a grade separation. It is expected that these walls could be rehabilitated

and placed back into service. A specific description of the problems observed at these five sites 

and their probable causes are provided in the paragraphs that follow.
143

The retaining walls at Américo Vespucio/Independencia (site 6) performed very well except for 

the detailing at the top of the modular block walls. The top two to three block layers and the cap 

block appeared to have been placed as a parapet above the ground surface at the wall top to 

hide a surface water drainage ditch (see figure 261). These blocks were unsupported and 

therefore only had self weight and the connection system to resist the horizontal loads. The 

horizontal force generated due to the earthquake and possibly from the concrete barrier behind 

the drainage ditch pushed the blocks and caused them to topple.

Figure 261. Photo. Toppling of coping at Américo Vespucio/Independencia.

The Chepe railroad bridge over the Biobío River (site 18) is a complex system, as shown in 

figure 176. A CIP cantilever wall supports a cut to form a depressed roadway. The ground water 

level is above the depressed roadway elevation at times, depending on the water level in the river.

In effect, the roadway and walls in this section form a tub to keep the ground water out of the 

roadway. The soil below and behind the wall settled and moved toward the river, probably due to 

liquefaction-induced slope failure (see figure 262 and figure 263). This condition is consistent 

with observations at other sites along this side of the Biobío River. Due to the liquefaction-induced

soil movement, the wall moved downward approximately 1 to 1.3 ft (0.3 to 0.4 m) and the top of 

wall rotated into the depressed roadway section. Furthermore, the seal between the roadway and 

wall was broken, which will allow water into the depressed roadway when the river rises.

Drainage 

ditch

Unsupported top 

block being held 

by connectors
144

Figure 262. Photo. Wall movement near Chepe railroad bridge due to 

lateral spreading and settlement.

Figure 263. Photo. Wall settlement and rotation due to lateral spreading and settlement.

While the wall was in immediate need of repair, it is notable that the wall was still performing its 

function of retaining the soil from impacting the depressed roadway despite the severe movement

caused by liquefaction. This demonstrates that even in the severe case of liquefaction below the 

wall, the wall can still function and not collapse. Note that the temporary support of the bridge 

superstructure placed after the earthquake may be contributing to the lateral loading of the wall.

This indicates that even though the wall was damaged, it has some reserve capacity.

The approach for the Juan Pablo II bridge (site 25) was discussed in section 5.1 and illustrated in 

figure 217 through figure 219. In addition to bridge damage, there were two CIP walls retaining 

the embankment fill between bridge spans (identified as north and south walls in figure 256, 

figure 257, figure 264, and figure 265). The embankment between bridge spans settled due to 

Settlement 

Lateral 

spreading

Wall Movement
145

liquefaction, and the bridge pier near the south wall also settled approximately 3 ft (1 m). The 

slopes in front of these walls moved laterally, likely due to liquefaction (see figure 200 and 

figure 201). It appears that the north wall and the embankment settled together more than 3 ft 

(1 m). Although the soil in front and behind the south wall settled, the wall itself did not settle 

because it was founded on deep foundations. This wall was built during a later construction 

project that added the entrance and exit ramps heading east. The north wall was completed in the 

original construction. Both walls continue to perform as intended after the earthquake, though 

the north wall would likely need to be replaced to raise the roadway back to its original grade.

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Figure 264. Map. Retaining walls near Juan Pablo II bridge.

Figure 265. Photo. Wall settlement near Juan Pablo II bridge due to 

liquefaction-induced settlement.

The retaining walls at Vía Elevada 21 de Mayo (site 28) are a combination of MSE walls used to 

retain approach and embankment fills and CIP concrete abutment and curtain walls. These walls 

were built in 1995 and are probably the first MSE walls built in Chile. These MSE walls were 

used as a substitute for CIP concrete walls included in the construction contract through a value 

engineering proposal made by the contractor prior to construction.

North 

South 

South North 

North
146

The MSE walls are composed of precast concrete panel facing and steel bar mat soil reinforcement.

There are a total of five walls on this site, and with the exception of the wall corners and full-height

joints between the MSE walls and the curtain walls, they performed very well.

Three wall locations within Vía Elevada 21 de Mayo experienced deformation and are highlighted

in this discussion (see figure 266). All three wall locations exhibited damage associated with 

wall top, full-height joint, and corner details. Corner details for MSE walls are an issue because 

they tend to attract seismic loads and ensuring that the joint will stay together can be difficult.

The performance issues were primarily with the joints opening up and fill escaping. Both bridges 

associated with the walls at this site also experienced damage but did not collapse. However, an 

older bridge parallel to these bridges did collapse.

© Google, GeoEye, DMapas, Inav/Geosistemas SRL, and Europa Technologies

Figure 266. Map. Vía Elevada 21 de Mayo.

Based on the design calculations available, the MSE walls at this site were designed using a 

horizontal acceleration coefficient of 0.4 g and were designed in general accordance with the

AASHTO Standard Specifications available at that time. The high acceleration coefficient used 

to design these walls was a bit surprising because the official implementation of the 0.4 g ground

acceleration for coastal areas did not occur until 2002. It is not known if the adjacent bridges 

were designed for 0.4 g. Based on the wall calculations, the soil reinforcement length used was

approximately 90–100 percent of the wall height, likely the result of using a 0.4 g acceleration 

coefficient. The corner and full-height joints where the walls connected to the bridge curtain walls 

were not designed as robustly for connectivity between adjacent panels as would typically be 

done today; this appears to have contributed to the performance problems observed.

Wall site 28A is a 30-ft (9-m) wall tiered perpendicular to the road alignment (see figure 267).

The wall appears to have rotated out, allowing backfill to escape out of the reinforced fill at the 

full-height joint between the curtain wall and the MSE wall. Potential issues could be related to 

North

Site 

Site 

Site 
147

inadequate soil reinforcement embedment lengths or poor placement and compaction of fill due 

to the tight geometry caused by the presence of the wall corner, abutments, and foundation 

elements. The uniform rounded nature of the sand backfill used (a coefficient of uniformity of 

less than 2 based on available laboratory test data on the backfill) may also have contributed to 

reduced pullout resistance, especially during shaking. Based on eye-witness accounts of the wall 

construction in 1995, the backfill soil was extremely difficult to compact to a firm condition 

because of the uniform nature of the soil. Although tiered walls are aesthetically pleasing and can 

help reduce stress on the lower portion of the wall, the use of such complex wall geometry can 

lead to poor performance if not designed, detailed, and constructed correctly.

 

Figure 267. Photo. Tiered wall corner tilting outward at wall site 28A.

Wall site 28B, a 33-ft (10-m)-high wall, is similar to wall site 28A in that the wall rotated 

outward, although not as severely (see figure 268). The cause of the problem is also similar to 

wall site 28A, a combination of inadequate soil reinforcement embedment due to the severe bridge

skew (the acute enclosed angle between the abutment wall and MSE wall is approximately 

25 degrees) and the use of uniform rounded sand. The combination potentially resulted in reduced

pullout resistance, especially near the wall top where soil confining stress is relatively low, and 

in lack of a positive connection at the full-height joint between the curtain wall and the MSE wall

panels. Modern technique to design such acute corners is to create a bin structure, making sure 

that the soil reinforcement is tied across to both sides of the acute angle. Since that was apparently

not done, the soil reinforcement near the acute corner was too short, allowing pullout failure to 

occur and the panels to move during the earthquake.
148

Figure 268. Photo. Wall corner tilting outward at wall site 28B.

The deformation at wall site 28C was the most severe of the three (see figure 269 through figure 272).

Similar to wall site A, the wall is tiered, but instead of rotating, the lower wall slid and rotated, 

moving laterally outward near the lower wall top approximately 1 to 1.3 ft (0.3 to 0.4 m) (see 

figure 269) and possibly 0.49 to 0.7 ft (0.15 to 0.2 m) near the wall base (estimated based on tilt 

of wall face). This is the only wall observed by TIRT that exhibited signs of translational movement,

with the exception of the wall at site 18 that moved laterally due to liquefaction-induced lateral 

spreading of the soil it retained. Because the upper tier was founded on the fill of the lower tier, 

the upper tier moved downward once the fill retained by the lower wall spilled out through the 

gap between the lower wall and the curtain wall. Figure 270 shows the back side of the upper 

wall, where the loss of the fill exposed the soil reinforcement behind the panels. The reinforcement,

panel connection, and lack of geotextile between the MSE wall and curtain wall can be seen. The 

function of the geotextile is to prevent fill from escaping through the joint. In this case, the gap 

was probably too large for the geotextile to keep the fill from spilling. The geotextile was used at 

the panel-to-panel joints. The movement of the corner panels was observed at the other corners 

of these walls, although to a lesser degree.

Bridge deck shifted 

5.9 to 7.8 inches 

(150 to 200 mm) 

toward acute skew 

angle

Top of wall 

moved out 

1ft (0.3 m) 
149

1 m = 3.28 ft

Figure 269. Photo. Lateral movement of lower wall top at wall site 28C.

Figure 270. Photo. Back side of upper wall at wall site 28C.

0.3 to 0.4 m
150

Figure 271. Photo. Fill material from tiered wall at wall site 28C.

Figure 272. Photo. Differential settlement at wall site 28C.

As shown in figure 271, the fill was a uniform medium-coarse sand with an angle of repose of 

approximately 32 degrees. This type of material may have contributed to poor soil-structure 

interaction with the reinforcement. If this was not considered in the wall design, it may have 

caused a lower pullout resistance than anticipated. 

Figure 272 shows the wall at the opposite abutment. In this case, the wall appears to have

differentially settled, causing the panels to shift and separate and allowing some backfill soil to 

spill out.

The specific cause of the excessive lateral movement and differential settlement at this location may

have been the liquefaction of the loose sandy soil below the wall. The available borings at the 

site indicate that the water table is approximately 7 ft (2 m) below the ground surface and loose, 

liquefiable sand extends to a depth of approximately 16 ft (5 m) (see appendix A). Soft organic silt 

extends another 3 to 7 ft (1 to 2 m) below the loose sand, below which the sands become dense to 

Panel 

gaps

32

Angle of repose
151

very dense. What appeared to be ejected sand (sand boils) was present near these walls as was 

some localized ground cracking and cracks in the roadway pavement, as shown in figure 273. With

regard to lateral movement on the lower wall, the width of the reinforced soil zone (approaching 

100 percent of the wall height) was likely more than adequate to resist the seismically induced 

lateral earth pressures behind the wall. Therefore, liquefaction is the most likely explanation for 

the lateral movement. During wall construction, 7.8 inches (200 mm) of settlement of the walls 

and fill occurred due to the loose sand and organic silt layer below the walls. Construction photos

of the walls show what appear to be gaps in the panels that were likely caused by differential

settlement that occurred during construction. However, it is also likely that liquefaction settlement

below the wall may have worsened the panel distortion.

Figure 273. Photo. Possible liquefaction features near wall site 28C.

The Muros Talca site (site 32) consists of four geogrid reinforced modular block MSE walls 

approximately 30 ft (9 m) high with a 1.5H:1V to 2H:1V unreinforced slope on top. The MSE 

walls abut to a CIP concrete abutment wall. A plan view of the bridge and walls is provided in 

figure 274. The northeast corner wall experienced severe cracking of the facing blocks and an

outward lateral deformation of approximately 4 inches (100 mm). The face block cracking pattern 

is shown in figure 275 and figure 276, with a linear distribution of the cracks on a 45-degree angle.

The southwest corner also had some cracked blocks with a similar pattern but to a much lesser 

degree. This appears to be another example of a wall corner (or very tight radius in the wall

alignment) having problems resisting seismic loading. The significant wall height, especially for 

a vertical wall, puts a lot of demand on the blocks even for static loading. Combining the wall 

height plus soil surcharge with a tight radius in the wall alignment and severe seismic loading 
152

caused a severe demand on the facing blocks. Some additional evaluation of this wall is needed 

to fully understand the effect that seismic loading had on this wall and the mode of facing failure.

However, even though the wall facing was sheared by the demands placed on it, the wall did not 

collapse and is still functional, though facing repair will be needed.

Figure 274. Illustration. Plan view of wall at Muros Talca.

1 mm = 0.039 inches

Figure 275. Photo. Wall with steep top slope at Muros Talca.

Observed 

cracked blocks

North

Wall face 

deformation 

100 mm max

100 to 150 

mm gap

Block cracking in 

45 degree shear bands
153

1 mm = 0.039 inches

Figure 276. Photo. Wall with cracked block pattern at Muros Talca.

Finally, it should be noted that the wall at Estribo Francisco Mostazal (site 11) is a geogrid reinforced

modular block wall directly supporting the abutment and bridge foundation loads (see table 6). This

type of abutment wall is what is referred to as a “true” MSE bridge abutment. The MSE abutment 

walls at site 11 performed very well, exhibiting no signs of lateral or vertical movement due to 

the earthquake. While the bridge suffered some relatively minor damage (see figure 277), the 

damage was not caused by the walls but was probably due to the severe bridge skew angle (similar 

to other bridges observed by TIRT and described in chapter 4) combined with the bridge tending 

to slide down hill, as it was located on a downhill roadway grade.

Figure 277. Photo. Bridge supported on footing on top of block-faced geogrid wall

at Estribo Francisco Mostazal.

Bridge railing 

offset

Bridge 

movement

Wall face 

deformation 

100 mm max

100 to 150 

mm gap

Block cracking in 

45 degree shear bands
154

In summary, the walls observed at the 32 sites visited by TIRT performed well, although a few 

walls did suffer some damage. In most cases, damage was minor and repairable. Of the walls that 

deformed laterally, movement tended to be greatest near the wall top with little, if any, lateral 

movement at the wall base, indicating that resistance to sliding is likely much greater than 

assumed in design. The most significant wall performance problems were primarily the result of 

inadequate details at the corners and full-height joints and at the wall top due to inadequate coping

details. These performance problems point to the importance of using details that ensure good 

connectivity between adjacent panels across a joint, which, for even current design codes such as 

the AASHTO LRFD bridge design specifications, should be better described. However, in spite 

of these performance problems, all of the walls met the no-collapse criterion typically applied for 

seismic design, including those walls that were subjected to the effects of liquefaction.
155

CHAPTER 6. LESSONS LEARNED FROM THE EARTHQUAKE

6.1 GENERAL

This chapter summarizes the overall performance of various bridge components, retaining walls, 

and bridge sites based on the observations detailed in chapter 4 and chapter 5. The observed 

performance leads to various lessons learned from the 2010 offshore Maule earthquake. It 

forms the basis for several conclusions drawn from the postearthquake reconnaissance regarding 

recommended seismic design and retrofit improvements and future research needs. 

6.2 SUPERSTRUCTURE ROTATION

6.2.1 Skewed Bridges

The direction of bridge skew is illustrated in figure 278 for clarity in the following discussions.

Skew is defined as the direction of rotation from the transverse line (perpendicular to the bridge 

centerline) to the skew side or abutment back wall of the bridge. For example, figure 278

indicates a clockwise skew and a counterclockwise skew. 

Figure 278. Illustration. Notations on skew direction.

Table 7 summarizes the characteristics and earthquake-induced damage patterns of bridges with 

a skew angle of 20 degrees or more. It lists the bridge name, bridge orientation, skew angle, 

skew direction, material used in girder, presence of diaphragms, transverse displacement at 

intermediate bents, and direction of deck rotation. The combination of these parameters helps

determine the mechanism of bridge deck rotations. 

Table 7. Summary of bridges and bridge damage.

Bridge Characteristics Damage Pattern

Site Name Orientation

Skew and

Direction

Girder/End

Diaphragm

Transverse 

Movement at

Intermediate Bents Rotation

1a

1b Miraflores NE-SW

20 degrees

counterclockwise Concrete/No Negligible Clockwise

2a Lo Echevers NE-SW

33 degrees 

counterclockwise Concrete/No Negligible Clockwise

5 Quilicura E-W

45 degrees 

counterclockwise Steel/Yes Negligible Clockwise

7 Romero E-W

31 degrees 

clockwise Concrete/No Negligible Counterclockwise

10b Hospital NW-SE

40 degrees 

counterclockwise Concrete/No Significant Clockwise
156

All of the bridges in table 7 consistently rotated about the centroid of the bridge superstructure in the

opposite direction of their skew, regardless of the bridge orientation, magnitude of the skew angle,

or presence of diaphragms. The fact that most of the bridges experienced negligible transverse 

displacements also indicates that the rotation effect was dominant in these bridge superstructures,

with the acute corners of each bridge moving away from their abutments at both ends. 

Based on the observations, the movement of a bridge superstructure can be illustrated in four steps,

as shown in figure 279. Under the earthquake excitations (steps 1a, 1b, and possibly 3), the bridge

superstructures first moved toward one abutment (left in figure 279) and impacted against the 

abutment back wall (step 2). The reaction from the back wall then turned the superstructure in a 

direction opposite to the skew direction (counterclockwise in figure 279). The rotational motion 

(step 3) was amplified due to the fact that the rotational vibration mode of the bridge superstructure 

is more sensitive to the ground motions, as illustrated with the acceleration response spectra 

recorded at Hospital Station in Curicó, since all the concrete girder bridges listed in table 7

have superstructures supported on neoprene pads and restrained with vertical seismic bars (see 

figure 280).

(6) The superstructures are weakly restrained in plan with the fundamental vibration 

mode in translation. With continuing deck rotations, the acute corners at two ends of the bridge 

finally moved away from the abutments, knocking off the curtain walls and becoming unseated 

(steps 4a and 4b). Note that the possibility of having significant rotational ground motions at the 

bridge site can further amplify the rotational motion (step 3).

Figure 279. Illustration. Deck rotation of a representative bridge (two spans shown).
157

Figure 280. Illustration. Rotational and translational mode responses to ground motions.

6.2.2 Straight Bridges

The bridges with little or no skew (Chada and Las Mercedes bridges) rotated counterclockwise,

as discussed in chapter 4. Unlike the skewed bridges, the deck rotation of the straight bridges 

cannot be explained by the skew effect. The possible factors contributing to significant rotations 

in these bridges are as follows:

The rotational mode of vibration of those bridges was very sensitive to ground motions. 

Any accidental eccentricity between the center of mass and the center of rigidity of the 

superstructure of a bridge could lead to substantial rotations.

The rotational component of ground motions could be significant.

The fault directivity effect could be significant because both bridges are approximately 

oriented along the east-west direction.

Considering both the skewed and straight bridges, observations can be made on the overall reasons

for bridge deck rotation. Skew of bridges is a significant but not necessarily decisive factor 

contributing to the bridge rotation. The high sensitivity of the rotational vibration mode of the

bridge to ground motion, particularly rotational excitations, could have been dominant. The fact 

that all bridges experiencing significant rotations were not far from Santiago, where soil conditions

are relatively stiff, supports the possibility of rotational ground motions at the bridge sites. Further

analysis is required to understand the significance of this phenomenon.
158

6.3 GIRDER DAMAGE

6.3.1 Fracture of Steel Girders

The superstructure of the Cardenal Raúl Silva Henríquez bridge is divided into two parts by the 

center expansion joints at bent 11. The northeast portion of the bridge is supported by a concrete 

substructure, and the southwest portion is mainly supported on a steel substructure. Most of the 

bridge bents are supported on drilled shafts. During the earthquake, the two parts most likely 

vibrated separately.

The most plausible reason for the girder damage at each abutment, as discussed in chapter 4, is 

excessive longitudinal force applied on the end support. During the earthquake, the majority of the

inertia force on half of the bridge superstructure was resisted by the end support at each abutment.

The excessive force resulted in either fillet weld fractures at the southwest abutment or steel girder

fractures in web and bottom flange at the northeast abutment.

6.3.2 Failure of Concrete Girders

The exterior prestressed concrete girders of the Chada and Romero bridges, which were constructed

without diaphragms, experienced out-of-plane block shear failures due to transverse impact loads 

from shear keys, as shown in figure 281. When partial diaphragms were used between the girders,

such as in the San Nicolás bridge, the bottom portion of the exterior girder still experienced

significant shear cracking, as shown in figure 282. However, with the use of even partial concrete 

teeth between girders, as shown in the Llacolen bridge in figure 283, both exterior and interior 

prestressed concrete girders suffered no visible damage. One interior girder in the west portion of 

Llacolen bridge experienced a horizontal crack, as shown in figure 57. This was because the 

concrete teeth provided sufficient lateral restraints on most of the girders, making them work 

together and share the transverse seismic force.

Figure 281. Photo. Exterior girder damage at Chada bridge, no diaphragms.
159

Figure 282. Photo. Exterior girder damage at San Nicolás bridge, partial diaphragms.

Figure 283. Photo. Exterior girder damage at west abutment of Llacolen bridge, 

concrete teeth.

6.4 CONNECTION BETWEEN SUPERSTRUCTURE AND SUBSTRUCTURE

6.4.1 Shear Key and Steel Stopper Failures

Concrete shear keys or shear keys in integral construction with curtain walls fulfilled the function

of sacrificial devices to protect the substructure of bridges. Their failures were observed regardless

of the presence of bridge deck rotations during the earthquake (e.g., Independencia, Chada, 

Romero, and Hospital bridges).

Steel stoppers used in several bridges (e.g., Independencia, Miraflores, and Lo Echevers bridges) 

failed prematurely. The Independencia bridge with steel stoppers was taken out of service following

the earthquake, but the parallel bridge with concrete shear keys and diaphragms survived the

earthquake with repairable damage. It was concluded that the two-bolt connections from each steel

stopper to the cap beam was too weak to resist any significant bending moment. However, once 

welded to steel girders, the steel stoppers prevented lateral movement of girders, as observed in 

the Quilicura railway bridge shown in figure 112.

6.4.2 Vertical Seismic Bars

Vertical seismic bars were used in a number of bridges (e.g., Chada, Las Mercedes, Llacolen, 

Romero, Hospital, and Pichibudis bridges). In general, they were flexible and underwent 

significant deformations during the earthquake. All were well anchored into the cap beams and 

decks. Field evidence only indicated bar pullout in bridges with fallen spans. It is uncertain 
160

whether the seismic bars provided vertical restraints to the bridge girders because there was no 

evidence of vertical up-and-down movements of the girders. This observation may be clarified 

after analyses using recorded ground motions.

6.4.3 Bridge Bearings

In general, bridge bearings functioned well during the earthquake. Several of them were displaced

significantly, such as in the Lo Echevers bridge shown in figure 95 and in the Cardenal Raúl 

Silva Henríquez bridge shown in figure 125. 

6.5 GIRDER SEAT LENGTH

The superstructure of a number of concrete and steel girder bridges (e.g., Llacolen, Miraflores, 

Lo Echevers, Romero, Hospital, Tubul, Biobío, and pedestrian bridges) dropped off their 

supports during the earthquake. In general, the support seat length is insufficient.

6.6 COLUMN SHEAR FAILURE

The pier caps and piers in all but the Llacolen and Juan Pablo II bridges received virtually no 

damage during the earthquake. This represented successful design of these bridge components. 

Several columns in the Llacolen and Juan Pablo II bridges failed in shear due mainly to the 

ground settlement and lateral spreading. 

6.7 FOUNDATION MOVEMENT AND DAMAGE

6.7.1 Overview

In general, bridge foundations performed relatively well in the earthquake. With the exception of 

cases where liquefaction-induced vertical or lateral soil movement was severe, foundations did 

not appear to suffer significant permanent deformations or significant damage, based on surficial 

observations. Most of the newer bridges visited were supported by shaft foundations (typically

4.9 ft (1.5 m) in diameter and less than 98 ft (30 m) deep). These foundations appeared to be

relatively light compared to bridge foundations currently used in areas of high seismic hazard in 

the United States. However, the foundations performed well in most cases.

To identify lessons learned regarding bridge foundations, two broad categories of geotechnical 

performance issues must be considered: foundation performance when liquefiable soils were 

probably not present and foundation performance when liquefiable soils were present. For those 

sites where liquefaction probably occurred, geotechnical performance issues are further divided 

between the effects of liquefaction-induced settlement and liquefaction-induced ground failure 

and the effects of lateral movement. The observations of geotechnical performance of foundations

in the sections that follow are made within this context.

6.7.2 Sites Not Affected by Liquefaction

Seventeen of the 32 sites visited did not appear to have significant liquefaction. However, poorer 

soil conditions (e.g., soft to still clays or loose sands) appeared to contribute to the amplification 

of ground motions (see chapter 2) and increased foundation deformation (see chapter 5), possibly 
161

increasing stress or deformation in the superstructure. In general, sites where better soil conditions

were present performed better and showed minimal damage. Near-surface soft to stiff clays 

appeared to be especially troublesome, especially in terms of their effect on approach fill stability

(e.g., sites 7, 8, and 10), but did not cause movement of or damage to the bridge abutments.

6.7.3 Sites Affected by Liquefaction

None of the sites where liquefaction likely occurred (15 of the 32 sites visited) were specifically 

designed to mitigate the effects of liquefaction through use of ground improvement or foundation 

strengthening (see chapter 5 for additional background on this issue). This affords the opportunity

to observe the effect liquefaction can have on foundation and abutment performance for structures

that are otherwise designed using the AASHTO or similar specifications, depending on the age 

of the structure.

6.7.3.1 Effects of Liquefaction-Induced Ground Failure

What was most surprising was the good performance of bridge abutments retaining 13- to 26-ft 

(4- to 8-m)-high approach fills over gently sloping ground, even when severe vertical and horizontal

approach fill deformation (1.6 to 3 ft (0.5 to1 m) or more) occurred due to liquefaction of soil below

the approach fill (see table 6, specifically sites 15, 20, 21, 22, and 27). Although there were a few 

cases where 1.9 to 5.8 inches (50 to 150 mm) of lateral movement of the abutment appeared to 

have occurred, in most cases, no discernable movement occurred. This may be the result of

three-dimensional effects reducing the lateral forces acting on the abutment foundations relative 

to what would be predicted assuming two-dimensional (i.e., plane strain) conditions. Furthermore, 

the liquefaction-induced slope failure tended to follow the path of least resistance (i.e., in the 

direction perpendicular to the roadway and bridge centerline). For those cases where either the 

abutment or an interior pier was located on a general slope, such as at a riverbank, the beneficial 

three-dimensional slope geometry was not present and foundation and substructure movement 

and damage due to liquefaction-induced ground failure was more likely (e.g., sites 17, 18, 19, 

and 25; see chapter 5 for details). These observations may have important implications for the 

strategy used for liquefaction design of bridges both in Chile and the United States.

6.7.3.2 Effects of Liquefaction-Induced Settlement and Downdrag

Liquefaction-induced ground settlement was observed at many of the bridge sites where liquefaction

occurred. However, settlement of the bridge foundations only occurred for a few of those sites 

(sites 25, 27, and 31). In general, regardless of the amount of liquefaction-induced ground settlement

that occurred, the foundations did not settle significantly if tipped in a reasonably good bearing 

layer. However, if the foundation was relatively shallow and not tipped in a relatively dense 

bearing layer, significant settlement of the foundation did occur.

6.8 RETAINING WALLS AND ROADWAY FILL

Three types of walls were inspected by TIRT: panel-faced MSE walls using bar mat or steel strip 

soil reinforcement, modular block high-density polyethylene geogrid reinforced walls, and concrete

gravity walls. Overall, retaining walls performed well during the earthquake. Tieback and soil nail

walls appeared to suffer little or no damage based on observations made by others. True MSE 

abutments, where the bridge footing foundation was directly supported on top of the MSE wall 
162

(e.g., site 11, see chapter 5), also performed well with no apparent deformation or damage to the 

walls. Note that most of these walls were designed using the AASHTO Standard Specifications (see

chapter 3). The observed wall performance appears to indicate that the AASHTO specifications 

as applied in Chile provide a safe design for seismic loading conditions.

There was no evidence of lateral sliding of the walls, the limit state that often controls wall design

for seismic conditions in North American design practice (see chapter 5). Where wall face lateral 

movement was observed, the movement was primarily rotational, with the maximum movement 

near the wall top. It appears that the passive resistance at the wall toe in combination with friction

along the wall base prevented significant translational movement.

Minor damage observed in several walls was due mainly to poor detailing. Inadequate coping 

details allowed a few of the top blocks to topple off the wall (modular block walls). Poor wall 

corner details or vertical full-height joint details (such as between the curtain wall and MSE wall 

retaining the approach fill sides) allowed panels to separate and wall backfill to spill out through 

the gaps in the facing. Stresses during seismic loading, especially for relatively tall walls (30 ft 

(9 m) or more) appeared to be more pronounced at abrupt changes in wall geometry (e.g., corners

and small radius changes in alignment), indicating the need for more robust wall facing designs 

in that type of situation. Soil reinforcement that is too short, especially near the wall top or where 

uniform low shear strength medium sand is used as backfill, can contribute to excessive wall or 

panel movement.
163

CHAPTER 7. CONCLUSIONS AND RECOMMENDATIONS

7.1 GENERAL

Based on the overall performance of various bridge components, retaining walls, and other 

structures summarized in chapter 6, conclusions can be drawn about the seismic performance 

of the transportation infrastructure in Chile during the offshore Maule earthquake. Considering 

the geologic conditions, earthquake settings, ground motions, and geotechnical and structural 

performance of the infrastructure, recommendations are made in this chapter for both immediate

implementation and future investigation. In addition, research needs are identified for the 

improvement of seismic design of bridges, walls, and other structures used in transportation

facilities. These recommendations may also be applicable for seismic retrofit purposes.

Some of the lessons learned require additional study to specifically assess how such lessons 

should be implemented. Toward that end, immediate needs are identified to gather more 

information and confirm assumptions and interpretations made in this report.

7.2 CONCLUSIONS ON TRANSPORTATION INFRASTRUCTURE PERFORMANCE 

Less than 0.15 percent of the bridges in the MOP inventory collapsed or suffered damage that 

rendered them unusable. Considering the magnitude of the earthquake, the transportation 

infrastructure in Chile performed relatively well and was able to support postearthquake 

emergency response and recovery. Even so, improvements can be made to further enhance 

the seismic performance of the infrastructure both in structural and geotechnical terms.

Many spans of precast prestressed discontinuous girder bridges with continuous decks fell from 

their supports, probably due to the significant in-plane rotation of their superstructures due to

severe shaking. Lateral steel stoppers used to provide both vertical and lateral restraint to the 

girders were largely unsuccessful due to their inadequate connection detail to cap beams or

abutments. Conversely, reinforced concrete shear keys served their design purpose, transferring 

lateral loads to the substructures. Vertical seismic bars were widely used to restrain the vertical 

motion of the spans, and they also performed well. Bridge substructures (foundation, column, 

and cap beam) generally behaved satisfactorily (undamaged cover concrete and no evidence of 

yielded reinforcing steel) except that several columns suffered shear failure due to ground

settlement and lateral spreading. However, the satisfactory performance of the columns generally 

occurred in bridges with failed superstructure-to-substructure connections, and it is not clear

whether this superior performance was due to overdesign of the columns or to the failure of the 

connections, which limited the forces in the columns to less than yield. All retaining walls 

exceeded performance expectations.

7.3 RECOMMENDATIONS FOR IMMEDIATE IMPLEMENTATION

7.3.1 Diaphragms in Concrete Girder Superstructures

In past earthquakes, full-depth diaphragms between girders have been very effective in distributing

earthquake loads among the girders and to the bearings and substructure below. Bridges with 

full-depth diaphragms again performed well in Chile, whereas those without diaphragms
164

performed poorly. To minimize the possibility of lateral shear failure and transverse unseating of 

girders, it is recommended that full-depth diaphragms be installed in all bridges and implemented 

in all future designs 

7.3.2 Support Length

Adequate support length is a proven countermeasure to the unseating of bridge spans. It is

recommended that generous seat lengths be provided in all new designs based on seismic hazard,

soil type, column height, distance between movement joints, and angle of skew. Both the AASHTO

specifications and Caltrans’ seismic design criteria have examples of such requirements. Existing 

bridges with inadequate support lengths should be retrofitted with seat extenders or longitudinal 

restrainers as a matter of priority 

7.3.3 Lateral Seismic Restraints 

Lateral steel stoppers of girders were typically anchored to their supports with two anchor bolts 

aligned in the direction of the span. In this case, only one bolt was available to transfer the

overturning moment when the seismic load from the girder was applied to the stopper some distance 

above the cap beam or abutment seat. This situation is quite ineffective. It is recommended that 

at least four anchor bolts be used to anchor each stopper to the cap beam in a configuration that 

provides both transverse and span-wise resistance to shear and overturning moments. 

Unless specifically designed to fail at lateral load that is less than the capacity of the column below

(thus protecting the column), all shear keys (both steel and concrete) should be designed against 

failure. This can be done in one of two ways: either the key or stopper is designed to resist the 

unreduced load from an elastic analysis (strength reduction factor = 1.0 or 0.8), or it is designed to

resist the maximum shear the column can generate in its fully yielded state (i.e., with fully developed

plastic hinges and including overstrength effects). Regardless of the method, the strength of any 

shear key should not be less than a specified minimum, usually expressed as a percentage of the 

tributary weight at the key. 

Shear keys are not necessary if the bearings are engineered to take the lateral loads and uplift 

forces. In such cases, the bearings need to have sufficient internal strength to transfer the forces 

without rupture and remain functional. They also need to be anchored accordingly. If it is

uneconomical to provide bearings of this type, shear keys are required. However the bearings 

should be anchored to the cap beams or abutments to prevent being dislodged during shaking. 

7.3.4 Substructure Design

Confinement has been demonstrated to be an effective means for preventing shear failure in 

reinforced concrete columns. It is recommended that columns with inadequate transverse 

reinforcing steel be retrofitted with external concrete, steel, or polymer jackets to prevent

brittle shear failures and to help ensure flexural ductile yielding. 

7.3.5 Retaining Wall Design 

Since retaining walls performed very well in general, new design specifications could be developed

to provide a no-analysis seismic design option for internal and external wall stability (i.e., for 
165

sliding, eccentricity, and bearing). Such an option might only be applicable within certain limits 

of ground acceleration, wall height, and surcharge conditions and would require minimum wall 

details (e.g., engineered connectivity of vertical joints in the face of the wall, especially at the 

corners; use of well-graded granular backfill; and good coping details on top of the wall).

The minimum wall details developed for the no-analysis option should also be required for walls 

that do not qualify for this option such as those in high seismic zones and should be considered 

for inclusion in the AASHTO specifications.

Bridges supported on MSE walls also appeared to have performed well in this earthquake. It is 

recommended that this cost-effective technology be investigated as an alternative to conventional, 

CIP, reinforced concrete abutment walls.

7.4 RECOMMENDATIONS FOR FUTURE WORK AND RESEARCH NEEDS

7.4.1 Structural Engineering Recommendations

Conduct a rigorous study of earthquake ground motion records to look for evidence of a 

rotational component in ground motion that could explain the in-plane rotation of 

non-skewed symmetric bridges. 

The ground motions recorded during the offshore Maule earthquake can be characterized by long 

duration and multiple pulses, particularly at those stations near Santiago where soils are relatively

stiff compared to other regions such as Constitución and Concepción. Although farther from the 

epicenter of the earthquake, the bridges near Santiago exclusively experienced significant in-plane 

rotations in their superstructures during ground shaking. The reason for this rotation is unclear,

but one possibility is that the stiffer soils near Santiago favored the propagation and amplification

of rotational ground motions.

To the authors’ knowledge, the occurrence of rotational components in earthquake ground motions

and their effect on bridge response and safety has not been previously studied. It is recommended 

that the ground motions recorded during this earthquake be rigorously analyzed and that case 

studies be undertaken for the non-skewed bridges that exhibited strong rotational response.

Study the relationship between the strength reduction factor, ductility, and period of vibration

during long-duration ground motions with multiple pulses. Also study the influence of these

characteristics in the ground motion on the in-plane rotation of non-skewed symmetric 

bridges. Prepare recommendations for AASHTO consideration regarding long-duration 

effects and potential changes to response modification factors.

It is known that long duration and multiple pulses can have a significant effect on the inelastic 

response of a bridge and can directly impact the integrity of the load path, especially if the bridge 

behaves as a pure elastoplastic system. In these circumstances, the relationship between the strength

reduction factor, ductility, and period of vibration may differ from that obtained under more

conventional ground motions and may warrant the use of a different set of response modification 

factors. The unique long-duration and multipulse feature of the ground motion may also affect the 

rotational response of bridge superstructures. These unique features in the ground motion and the 
166

corresponding inelastic response in the structure and approach fills may also explain the in-plane 

rotation of the superstructures in non-skewed bridges. 

Considering the similarities in geologic and tectonic conditions, design specifications, and 

construction practices between the Chile and the United States, indepth studies on the effect of 

long-duration ground motions (with and without multiple pulses) on bridge behavior during this 

earthquake would be of immediate relevance to U.S. practice, particularly in Oregon and 

Washington. Research outcomes are likely to affect ground motion provisions as well as design 

requirements for bridges in these states.

Validate the support length requirements in the AASHTO specifications for skewed bridges. 

Many spans of skewed bridges were unseated, confirming the sensitivity of this kind of bridge to 

collapse. The minimum requirements in the AASHTO specifications are based on engineering 

judgment rather than rigorous analysis, and this earthquake provides the opportunity to check the 

validity of these requirements, particularly for skew. Research outcomes would likely affect the 

support length requirements for bridges throughout the United States.

Study the implications of combining the scour and earthquake load cases for coastal bridges

located in tsunami inundation zones

The earthquake-induced tsunami generated significant waves and currents that added both horizontal

and vertical loads on some coastal bridges and caused the erosion of soils around bridge foundations. 

These additional loads may further damage a bridge that has already been damaged by earthquake

shaking. Structures may collapse and low-lying bridges may be lifted off their substructures and 

swept upstream. Although designing for scour is required in U.S. specifications, the combination 

of the effect of scour and earthquake effects is not required, as in the Chilean code (see appendix B).

These combined effects warrant further investigation 

Conduct feasibility studies of bridges without diaphragms for potential application in 

accelerated bridge construction.

Concrete blocks (or shear keys) on each side of cap beam girders are an alternative to diaphragms

for the distribution of transverse loads between girders and from the deck slab to the cap beam.

They can effectively prevent the girders from shear failure by reducing the maximum load on 

each individual girder. If the joint between the top flange of the girder and deck slab can be 

economically designed to transfer significant moments without distress, this alternative may be 

an attractive strategy in accelerated bridge construction 

Conduct case studies on the AASHTO global design strategy that permits fusing of the 

super-to-substructure connection and protection of the substructures and foundations from 

seismic damage.

Adjusting the strength of the shear keys in a bridge superstructure is an alternative strategy for

managing the load path and energy dissipation, thus enabling cost-effective designs that ensure 

bridge substructures suffer little or no damage. This strategy is permitted in the new AASHTO 

specifications for seismic design, but no comprehensive study has been done to date to prove its 

viability. This earthquake provides the opportunity to fill this gap.
167

Study the trade-off between lower maintenance costs with fewer joints and reduced 

redundancy for longitudinal seismic loads in continuous bridges.

Jointless bridges are widely favored for their lower corrosion-related maintenance costs and the 

increase in redundancy for gravity loads. However, the longitudinal distribution of load in these 

continuous bridges can be a challenge, especially in those cases where the superstructures are not 

monolithic with their piers (i.e., supported on bearings that allow thermal expansion). In long, 

multispan bridges, very large loads can be attracted into a few structural elements, and, unless 

designed for this heavy demand, failure can occur in the super-to-substructure connections at 

these elements. Studies are required to understand the trade-off between lower maintenance costs 

(with fewer joints) and the increased demand on a few substructures (i.e., the reduced 

redundancy for lateral loads), 

7.4.2 Geotechnical Engineering Recommendations

As previously noted, geotechnical conclusions in this report are based on surficial observations 

and approximate geometry. Subsurface data and foundation and geometric details are required 

for most of the sites visited to confirm these preliminary interpretations. These data are a necessary

first step towards deriving the most benefit from the investigation of this earthquake. Digitized 

ground motion records are also required if the more detailed modeling necessary to confirm the 

preliminary conclusions is to be achieved. Once site-specific data are obtained, future efforts in 

the geotechnical area should focus on the following:

Develop the ground motion records from this earthquake for use in seismic design when 

long-duration subduction zone records are needed and add them to databases used by 

designers for conducting site-specific seismic analyses.

Achieving this goal will involve obtaining subsurface information and other site information 

where the ground motions were recorded so that they can be deconvolved to base rock motions, 

cleaned up, baselined, etc. in order to make them most useable. Such digital records would also 

be extremely useful for design where subduction zone earthquakes must be considered (e.g., the 

west coast of North America).

Conduct research to better quantify the beneficial effect of three-dimensional geometry 

issues on abutment performance when liquefaction occurs.

For example, how wide does the fill need to be to eliminate the three-dimensional benefit identified

in chapter 5 and chapter 6 such that the use of a two-dimensional slope stability or lateral 

spreading analysis in the direction toward the bridge is not too conservative? The case histories 

summarized in this report should be used to calibrate three-dimensional numerical models to 

investigate this issue more fully.

Conduct research to improve the quantification of lateral and vertical (downdrag) forces 

on foundations and abutments induced by liquefaction.

Based on the foundation performance observed in this earthquake, liquefaction lateral spreading 

and downdrag design with regard to their effect on foundations may be too conservative. 
168

Conduct research to improve the estimation of lateral and vertical ground deformation 

caused by liquefaction using measured deformation data from the 15 bridge sites where 

liquefaction occurred as a result of this earthquake.

Methods currently available for estimating these deformations are relatively crude. In addition, a 

reasonable estimate of both vertical and horizontal ground movements due to ground liquefaction 

and lateral spreading should be made available to practitioners who will implement the displacementbased design methodology to ensure that undesirable structural failure and damage modes of bridges

are removed. The fact that most liquefied sites during the Chile earthquake were not treated for 

improved performance offers a good opportunity to improve the engineering estimate of 

liquefaction-induced vertical and horizontal ground movements.

Use the liquefaction performance research identified above to develop a strategy from a nocollapse design objective perspective to focus liquefaction mitigation efforts on only the 

more severe cases.

Judging what should be considered severe may be a challenge and may be very site-specific.

Even if such a strategy could be formulated and implemented, thorough site- and project-specific 

geotechnical analyses would still be needed to assess the severity of the likely effects of

liquefaction at a given site.

Conduct research to develop more accurate lateral deformation models for walls under 

seismic loading considering the likelihood of rotational rather than translational movement.

Since very little evidence of the translational movement of a wall as a rigid body was found, the 

theoretical approach typically used to perform seismic deformational analysis of walls (Newmark

analysis) may not match reality.(14,20) However, the good seismic wall performance observed still 

provides justification for using a reduced horizontal acceleration coefficient for design, even 

considering the lack of translational movement. 

Conduct research to investigate the effect of high vertical acceleration on the performance 

of selected bridges and walls.

Vertical accelerations are often ignored in current seismic design practice. The effects of these 

relatively high vertical accelerations on bridge and wall performance should be further 

investigated and their implications to design practice determined.
169

APPENDIX A. SUBSURFACE DATA

A.1 SUBSURFACE DATA

A.1.1 Subsurface Data for Américo Vespucio/Miraflores (Site 1)

One boring log was available at site 1. The specific location is not known, but its general 

location is within the circle in figure 284. A copy of the boring log is provided in figure 285.

© Google, Europa Technologies, Inav/Geosistemas SRL, and GeoEye

Figure 284. Map. Image of site 1 taken prior to earthquake.

Source: MOP

Figure 285. Illustration. Boring log at site 1.

Silty Clay

Layers of 

silty Clay 

and fine 

sand 

Silty fine 

sand 

Clay

Hard Silt

Clay and 

silty Clay
170

A.1.2 Subsurface data for Américo Vespucio/Lo Echevers (Site 2)

The location of the site 2 bridge is shown in figure 286. A copy of as-built bridge plans and one 

boring log at site 2 are provided in figure 287 through figure 289.

© Google, Europa Technologies, Inav/Geosistemas SRL, and GeoEye

Figure 286. Map. Image of site 2 taken prior to earthquake and showing 

approximate location of test hole.

Source: MOP

Figure 287. Illustration. As-built bridge plan for site 2.

Test Hole
171

Source: MOP

Figure 288. Illustration. As-built bridge profile for site 2 showing footing foundations.

Source: MOP

Figure 289. Illustration. Boring log at site 2.

A.1.3 Subsurface Data for Avenida Romero Accesso Sur (Site 7)

The location of bridge site 7 is shown in figure 290. A copy of as-built drawings and one boring 

log at site 7 are included in figure 291 and figure 292.

Clay

Silty clay, 

silty sand,

silty clay, 

and fine 

sand

Layered 

silt and 

clay
172

© Google, DigitalGlobe, and Inav/Geosistemas SRL

Figure 290. Map. Image of site 7 taken prior to earthquake and showing 

approximate location of test hole.

Source: MOP

Figure 291. Illustration. As-built bridge plan for site 7.

Test Hole S-2
173

Source: MOP

Figure 292. Illustration. Boring log for site 7.

A.1.4 Subsurface Data for Avenida Chada Accesso Sur (Site 8)

The location of bridge site 8 is figure 293. A copy of as-built drawings and one boring log at 

site 8 are included in figure 294 and figure 295.

© Google, DigitalGlobe, and Inav/Geosistemas SRL

Figure 293 Map. Image of site 8 taken prior to earthquake and showing 

approximate location of test hole.

Fine soil

Sand

Gravel

Fine soil

Sand

Fine soil

Test Hole SA-3
174

Source: MOP

Figure 294. Illustration. As-built bridge plan for site 8.
175

Source: MOP

Figure 295. Illustration. Boring log for site 8.

See ASTM 

D2488 for 

definitions of soil 

classifications
176

A.1.5 Subsurface Data for Route 5 Railway Overcrossing at Hospital (Site 10)

The location of bridge site 10 is shown in figure 296. A copy of as-built drawings and one boring 

log at site 10 are included in figure 297 and figure 298.

© Google, DigitalGlobe, and Inav/Geosistemas SRL

Figure 296. Map. Image of site 10 taken prior to earthquake and showing 

approximate location of test hole.

Source: MOP

Figure 297. Illustration. As-built bridge plan for site 10.

Test Hole SA-2
177

Source: MOP

Figure 298. Illustration. Boring log SA-2 for site 10.

A.1.6 Subsurface Data for Via Elevada 21 de Mayo/Cruce Ferroviario (Site 28)

The location of bridge site 28 is shown in figure 299. Red dots indicate the location of the test 

holes with the corresponding numbers labeled. Copies of three boring logs at site 28 are given in 

figure 300 through figure 305.

© Google, DigitalGlobe, Inav/Geosistemas SRL, and Europa Technologies

Figure 299. Map. Image of site 28 taken prior to earthquake and showing 

approximate locations of test holes.

See ASTM 

D2488 for 

definitions of soil 

classifications

S2

S1

S3

Site B

Site A

Site C
178

Source: W. Neely

Figure 300. Illustration. Boring log S-1 for site 28.

Coarse 

sand

Organic 

silt

Silty fine 

sand

Water table 

2-12-91

Fine 

S d

Silty fine 

Sand

Organic 

Silt

Medium 

Sand
179

N1 = number of blows per 15 cm.

N2 = number of blows per 30 cm.

Entre colas = depth range.

1 cm = 0.39 inches.

Source: W. Neely

Figure 301. Illustration. Detailed standard penetration test blow counts 

for boring log S-1 for site 28.

Boring No. 1
180

Source: W. Neely

Figure 302. Illustration. Boring log S-2 for site 28.

Silty coarse 

Sand

Water table 

3-11-91

Plastic Clay

Fine Sand

Plastic Silt

Silty with 

fine Sand

Fine Sand

Organic Silt

Plastic silty 

fine sand

Fine Sand
181

N1 =number of blows per 15 cm.

N2 =number of blows per 30 cm.

Entre colas = depth range.

1 cm = 0.39 inches.

Source: W. Neely

Figure 303. Illustration. Detailed standard penetration test blow counts 

for boring log S-2 for site 28

Boring No. 2
182

Source: W. Neely

Figure 304. Illustration. Boring log S-3 for site 28.

Fine to medium 

Sand

Water table

4-12-91

Organic Silt

Silty fine Sand

Medium Sand 

with silt

Plastic Silt

Fine Sand 
183

N1 =number of blows per 15 cm.

N2 =number of blows per 30 cm.

Entre colas = depth range.

1 cm = 0.39 inches.

Source: W. Neely

Figure 305. Illustration. Detailed standard penetration test blow counts 

for boring log S-3 for site 28.

Boring No. 3

185

APPENDIX B. SEISMIC DESIGN REQUIREMENTS FOR BRIDGES AND WALLS 

IN CHILE

B.1 SEISMIC DESIGN REQUIREMENTS FOR BRIDGES

Section 3.1004 of the MOP Highway Handbook establishes the seismic design requirements for 

small to medium span bridges (viaducts, overheads, and pedestrian bridges) with span lengths 

less than 230 ft (70 m ).(11) Longer bridges and those with special structural configurations

(arches, suspension, and cable-stayed bridges) are not covered by these specifications.

The philosophy underlying these specifications is to ensure public safety, but significant damage 

and disruption to service is accepted. The ground motions and design forces are calculated based 

on a probability of exceedance of 10 percent in 50 years, which is equivalent to a return period of 

475 years. For design purposes, the country is divided into nine regions from north to south and 

in three seismic zones from east to west. These zones are determined by the value of the PGA in 

stiff soil, as shown in table 8. Zonation maps are included in the seismic provisions for all of the 

nine regions. The map for central Chile is shown in figure 22.

Table 8. Seismic zones.

Seismic 

Zone

PGA, 

Ao

Seismic 

Coefficient

1 0.20 g 0.20

2 0.30 g 0.30

3 0.40 g 0.40

The effect of the soil stiffness on the values of PGA at a specific site are taken in account by 

multiplying the values in table 8 by a soil coefficient defined in table 9.

Table 9. Soil coefficients.

Soil 

Type Description S

I Rock with measured shear wave velocity, Vs > 2,600 ft/s (800 m/s) 0.9

II

Very dense soil with measured shear wave velocity, 1,300 ft/s (400 m/s) < Vs

< 2,600 ft/s (800 m/s) 1.0

III Soft to medium-stiff clays and sands. Unsaturated sands with N > 20 blows/ft 1.2

IV Saturated cohesive soils with Su > 0.0036 ksi (0.025 MPa) 1.3

S = Soil coefficient.

Vs = Shear wave velocity. 

N = Blow count number.

Su = Average undrained shear strength in top 100 ft (30 m).

For seismic zones 2 and 3, an importance coefficient is included in the calculation for seismic 

forces. Bridges are classified as either “essential” (importance coefficient I) or “other” 

(importance coefficient II).
186

Since scour around or under a bridge foundation may influence its seismic behavior, the 

consequences of this hazard are included in the design. Two values of a seismic scour index (PSS)

are defined. A value of PSS equal to 1 requires 75 percent of the total depths of scour to be

considered in design, and a value of 2 requires 100 percent of the total scour to be considered in 

design. The depth of scour is included in the determination of the height of the superstructure 

(substructure slenderness) and consequently affects the appropriate analysis method.

Four categories of seismic behavior are defined based on the PGA, PSS, and the importance 

coefficient, as shown in table 12.

Table 10. Seismic behavior categories.

PGA, Ao PSS

Importance 

Coefficient

I II

0.20 g

1 a a

2 b b

0.30 g

1 b b

2 c c

0.40 g

1 c c

2 d d

B.1.1 Analysis Methods

One of five analysis methods is used in to determine the elastic seismic forces, depending on the 

geometry and importance of the bridge. These methods are: the seismic coefficient method, the 

seismic coefficient method modified by structural response, the modal spectral analysis method, 

the modal spectral analysis method using site-specific spectra, and the time history analysis 

method (linear and nonlinear).

B.1.1.1 Seismic Coefficient Method

This method is applied to simply supported bridges with a maximum of two spans (each less than 

230 ft (70 m) and substructure heights less than 39 ft (12 m). The substructure height is measured

between the seat and the foundation soil including the maximum scour effect. The horizontal 

seismic coefficient Kh is calculated as shown in equation 1.

(1)

where:

K1 = importance factor defined in table 11.

S = soil coefficient defined in table 9.

Ao = peak ground acceleration defined in table 8.

g = acceleration due to gravity (32 ft/s2 (9.8 m/s2)).

Kh=K1. S .

Ao

2g

≤0.10
187

Table 11. Importance factor, K1

Importance 

coefficient K1

I 1.0

II 0.8

In this method, all structural components are designed to be elastic, which means the response 

modification factor is equal to one. The analysis is done for both the longitudinal and transverse 

directions.

B.1.1.2 Seismic Coefficient Method Modified by Structural Response

This method is applied to simply supported bridges of more than two spans (spans lengths less

than 230 ft (70 m)) and substructure heights less than 82 ft (25 m) including the maximum scour 

effect. The horizontal seismic coefficient Kh is calculated as shown in equation 2.

(2)

where: 

K2,T1 = spectral constants, defined in table 12.

Tn = natural period determined using simplified formulas (table 13) or exact methods.

Vb = total base shear in the bridge.

P = total weight of the bridge.

Table 12. Spectral constants, T1 and K2.

Soil type T1 K2

I 0.20 0.513

II 0.30 0.672

III 0.70 1.182

IV 1.10 1.598

Kh(Tn)=

⎩

⎪

⎨

⎪

⎧1.5 K1S Ao

g

Tn≤ T1

K1K2S Ao

gTn

2/3 Tn> T1

Vb≥0.22 K1. S .

Ao

2g . P
188

Table 13. Approximate formulas for natural periods for bridges.

Structural system Direction Fundamental period

Simply supported 

bridges

Longitudinal or 

transverse

Reinforced 

concrete 

support

Continuous bridges 

with fix supports on 

intermediate 

substructure and 

rigid abutments. 

One superstructure 

end on a fix support

Transverse1 Reinforced 

concrete 

Steel

Longitudinal

1 This formula should not be applied to multiple column bents.

Wp = weight of the pier or column (ton).

Wu = weight of the superstructure supported on the pier or column (ton).

Ec =elastic modulus (ton/m2

).

I = moment of inertia of the file or column in the direction of analysis (m4

).

H = height of the pile or column, measured as the difference in levels between the seat and the foundation soil 

considering the maximum scour effect (m).

B.1.1.3 Modal Spectral Method

This method is applied to tall bridges with simple supported continuous bridges (spans lengths less

than 230 ft (70 m)) and substructure heights greater than 82 ft (25 m) including the maximum 

scour effect. This method may be used in place of the seismic coefficient method and the seismic 

coefficient method modified by structural response.

The absolute spectral acceleration for the vibration mode m, Sa(Tm) is calculated using equation 3.

Vb ≥ 0.20 K1. S .A0. P/g

(3)

where: 

Tm = period of mode m.

The displacements, rotations, and forces in each structural element are calculated in both the 

longitudinal and transverse directions from the contributions of each of the vibration modes. The 

combination of the maximum modal values to obtain elastic response values are calculated using 

the expressions shown in equation 4 and equation 5.

Tn=2π �

0.3 Wp+Wu

3 EcI g H3

Tn=2π �

0.3 Wp+Wu

3 EcI g H3

Tn=2π �0.3 Wp+Wu

4.5 EcI g H3

Tn= π

8 � Wp

EcI gH3

Sa(Tm)= �

1.5 K1S Ao Tm≤ T1

K1K2S Ao

Tm

2/3 Tm> T1
189

(4)

(5)

where:

= summation of the contribution of the considered modes.

= coupling factor.

Ti, Tj = modal periods.

= damping ratio and taken equal to 0.05.

The number of modes to be included in the analysis is chosen such that the summation of the 

effective mass ratios will be larger than 90 percent.

B.1.1.4 Modal Spectral Method Using Site-Specific Spectra

In special cases, MOP requires that the design of a bridge is based on the modal spectral method 

of analysis based and a spectrum developed for the specific site.

B.1.1.5 Linear and Nonlinear Time History Analysis

In specific cases, especially for bridges incorporating base isolation systems, MOP requires time

history analysis, either linear or nonlinear, using synthetic accelerograms reflecting the site 

seismic risk and the soil conditions. 

B.1.2 Response Modification Factors (R)

The design seismic forces for members and connections are calculated by dividing the elastic 

response values by a response modification factor R. Special detailing is required to accommodate

inelastic behavior and the formation of plastic hinges when R factors greater than 1 are used.

Values of these factors are summarized in table 14 for the longitudinal and transverse directions. 

S=���ρij Si

j

Sj

i

ρij= 8 ξ

2

r

3/2

(1+r)(1-r)2+4 ξ

2

r (1+r)

 r=

Ti

Tj

≤1.0

ΣiΣj

ρij
190

Table 14. Response modification factors (R).

Element Longitudinal Transverse

Supports

Wall-type pier 3 2

Single column with foundation cap 3 3

Column bent with foundation cap 3 4

Inclined column with foundation cap 3 2

Piles or micropiles

Individual 3 3

Batter piles 3 4

Inclined batter piles 3 2

Foundations1

Spread footing 1 1

Batter of piles 1 1

Drilled shaft 1 1

Caisson or foundation pile 1 1

Connections2

Expansion joint 0.8 0.8

Bearing plate 0.8 0.8

Shear key 1 1

Bearing 1 1 1 For bridges in seismic categories C or D, seismic forces are those developed by plastic 

hinges in the column. These forces are frequently less than those calculated using R = 1. 2 Connections are all elements that transfer shear or axial loads from one point to another in 

the structure. Moment connections are not included in this definition.

B.1.3 Determination of Elastic Forces and Displacements 

For bridges classified in seismic categories C or D, the elastic forces and displacement are 

calculated in both the longitudinal and transverse directions. For the case of curved bridges, the 

longitudinal direction is the chord joining the abutments.

B.1.4 Combination of Forces in Two Perpendicular Directions

The forces calculated independently in two perpendicular directions are combined to take in 

account the uncertainty in the direction of the seismic waves and the effect of the two horizontal 

components of the earthquake. Two load stages are considered. Load stage I is shown in equation 6,

and load stage II is shown in equation 7. 

(6)

(7)

where:

F = absolute value of force or moment for all structural members due to load combination.

FL = absolute value of force or moment in the longitudinal direction.

FT = absolute value of force or moment in the transverse direction.

F = 100% FL + 30% FT

F = 30% FL + 100% FT
191

B.1.5 Minimum Support Length

Minimum values for support lengths of beams are shown in equation 8 for seismic behavior 

categories A and B and in equation 9 for seismic behavior categories C and D. 

(8)

(9)

where:

L = length of the bridge deck to the adjacent expansion joint, or to the end of the bridge deck; for 

hinges within a span, L is the sum of the distances to either side of the hinge; for single-span 

bridges, L is the length of the bridge deck (m).

= angle of skew of support measured from a line normal to span (degrees).

H = 0 for abutments and single-span bridges; average height of columns supporting the bridge 

deck from the abutment to the next expansion joint (m); pier height (m) for columns and 

piers; average height of the adjacent two columns or piers (m) at hinges within a span.

B.1.6 Modified Design Forces for Structural Elements and Connections

The seismic design forces from load stages I and II are divided by the R factors to obtain

modified design forces. These modified forces are combined with other loads to obtain the 

maximum load combination as shown in equation 10.

(10)

where:

D = dead load.

B = buoyancy.

SF = stream-flow pressure.

E = earth pressure.

EQM = elastic seismic forces for the load stages I and II divided by R.

B.1.7 Modified Design Forces for Foundations

The seismic design forces for spread footings, pile caps and micropiles shall be those from load 

stages I and II divided by the response modification factor R = 1 to obtain modified design 

forces. These modified forces are combined with other loads to obtain the maximum load 

combination as previously shown in equation 10.

The load combinations are based on division I-A of AASHTO’s Standard Specifications.(10)

N=(203+1.67L+6.66H)(1+0.000125 α2) (mm)

N=(305+2.5L+10H)(1+0.000125 α2) (mm)

1.0(D + B + SF + E + EQM)
192

B.1.8 Seismic Design of Foundations, Abutments, and Retaining Walls

The specifications for the seismic design of these elements are covered in section 3.1003 of the 

Highway Handbook.(11)

B.1.9 Seismic Design of Reinforced Concrete Elements

The design and construction of CIP reinforced concrete elements (columns, foundations, and 

connections) are required to satisfy the provisions of division I-A of the AASHTO specifications.

If the design is based on the allowable stress method, all the stresses are increased by 33.3 percent.

If this method is used for bridges in seismic categories C and D, conservative designs are to be 

expected unless the design is based on the maximum forces developed from plastic hinging.

B.1.9.1 Requirements for Columns

Section 7.6.2 of division I-A of the AASHTO specifications for seismic performance category C 

and D is followed.

B.1.9.2 Columns and/or Piers

Section 7.6.3 of division I-A of the AASHTO specifications for seismic performance category C 

and D is followed.

B.1.9.3 Column Connections

Section 7.6.4 of division I-A of the AASHTO specifications for seismic performance category C 

and D is followed.

B.1.9.4 Construction Joints in Piers and Columns

Section 7.6.5 of division I-A of the AASHTO specifications for seismic performance category C 

and D is followed.

B.1.9.5 Reinforced Concrete Micropiles

These elements are designed for elastic forces (R = 1). Alternatively, these elements may be 

designed using the maximum forces from the column or pier plastic hinges.

B.1.10 Forces from Plastic Hinges on Piers or Columns

Design forces for columns, piers, and foundations are determined from the maximum forces due 

to plastic hinges. The procedure for calculation of these forces is based on section 7.2.2 of 

division I-A of the AASHTO specifications.

B.1.11 Diaphragms

Due to the high values of vertical acceleration in seismic zone 3, the beams of bridges in this 

zone are connected by transverse beams (diaphragms).
193

B.1.12 Anchor Bars (Seismic Bars)

No.7 (22 mm) deformed bars are provided to resist a vertical seismic coefficient of Kv = Ao/4g.

These bars are required to satisfy the provisions of ASTM A706. ASTM A 615M bars are permitted 

if the difference between the true yield force and the nominal value is less than 17.4 ksi (120 MPa)

and the relationship between ultimate stress and yield stress is less than 1.25.

B.1.13 Lateral Stoppers (Shear Keys)

To avoid excessive lateral displacement of the superstructure due to seismic forces, lateral stoppers

are provided on bent caps. The design force used for these stoppers is the seismic force acting in 

the transverse direction divided by four. These stoppers are required to have sufficient ductility 

to avoid the deck collapse. The height of these shear keys is required to be greater than 12 inches 

(300 mm) and a gap between the shear key and the superstructure is provided to accommodate

the seismic displacement plus 2 inches (50 mm). 

B.1.14 Seismic Joints

Seismic joints between deck segments have a minimum spacing to allow the displacements of 

the elastomeric bearings. The minimum spacing is given by equation 11:

(11)

where S1 and S2 are the seismic displacements of each of the elastomeric bearings in cm.

B.1.15 Seismic Isolation Systems

Seismic isolation systems used as supporting elements are designed and tested according to 

AASHTO’s Guide Specifications for Seismic Isolation Design.

(21)

B.2 SEISMIC DESIGN REQUIREMENTS FOR WALLS

Walls that are part of the public transportation system controlled by MOP are generally designed 

in accordance with the AASHTO Standard Specifications.(10) PGAs given in the Chilean design 

guide for wall design are used (i.e., PGAs of 0.2, 0.3, or 0.4 g in zones 1, 2, or 3, respectively). 

MOP requirements allow concrete gravity walls to be designed for a reduced acceleration,

providing that the wall can slide during shaking. For MSE walls however, MOP requirements do 

not allow the design acceleration to be reduced for internal and external wall stability. For walls 

not under MOP jurisdiction, design accelerations may be established as a site-specific value. In 

general, however, even these walls are designed using the AASHTO Standard Specifications.(10)

For MSE walls, typical practice is to select good, frictional backfill materials and overexcavate 

poorer soils, replacing them with the backfill materials. Because of this practice, a reinforcement 

length of 70 percent of the wall height is typically used to meet seismic design requirements.

Sj≥6.25 

Ao

g

+S1+S2 (cm)

195

ACKNOWLEDGMENTS

Financial support for the postearthquake reconnaissance trip to Chile was provided mainly by 

FHWA, with Sheila Duwadi serving as the point of contact for U.S. team members. The authors

would also like to acknowledge the financial support of EERI. The findings and opinions expressed

in this report are those of the authors and do not necessarily reflect those of the sponsors. Special 

thanks are due to José Miguel Ortega for his support and coordination with local authorities to 

make the reconnaissance trip a success, Mauricio Guzmán and Sandra Achurra for their escort 

and introduction to various bridge sites, Alex Unión Villarroel for his posttrip assistance, and 

Rodrigo Oviedo for his participation throughout the reconnaissance trip. 

197

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HRDI-50/10-11(200)E