Study on Post 

Tensioning Methods

Office of Nuclear Regulatory Research

NUREG/CR-7208
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Study on Post 

Tensioning Methods

Manuscript Completed: November 2015

Date Published: November 2015

Prepared by: 

Christopher A. Jones and Lili A. A. Heitman 

Sandia National Laboratories

1515 Eubank

Albuquerque, NM 87123

Robert Dameron

Moffatt and Nichol

1660 Hotel Circle North

San Diego, CA 92108

Madhumita Sircar, NRC Project Manager and Technical Lead

NRC Job Code V6186-00

Office of Nuclear Regulatory Research

NUREG/CR-7208

 

iii

ABSTRACT

In order to assess the effects of using grouted prestressing tendon systems in nuclear power 

containment vessels, a containment vessel is modeled in three dimensions using finite element 

analysis and different assumptions relating to grouted and ungrouted prestressing systems are 

applied. The particular containment vessel is the 1:4 Scale Prestressed Concrete Containment 

Vessel, for which extensive data exists. Such data can be used to benchmark and validate the 

modeling. The results from the grouted and ungrouted models indicate that greater localized 

maximum stresses and strains are predicted for the grouted case because the tendon system is 

not permitted to slip and redistribute loads as the vessel deforms. Accordingly it is noted that the 

analyses predict failure of the vessel at a slightly lower internal pressure in the case of the 

grouted system. Additionally, the response of the containment structure to tendon corrosion is 

investigated for both grouted and ungrouted systems while the prestressing system is subject to 

corrosion. Contrary to the results from the uncorroded tendon simulations, when corrosion 

impacts the prestressing system, the grouted system offers better performance owing to 

isolation of the corrosive effects and better preservation of prestressing force. 

A post-tensioned containment is not only subject to normal material degradation, but also to a 

continuous loss of prestressing force. The losses result from inherent material characteristics. 

Because loss of prestress is not completely predictable, loss of prestress or residual strength 

must be measured at regular intervals to determine the containment strength to resist accident 

pressure and design loads with acceptable margins. The US Nuclear Regulatory Commission 

and Sandia National Laboratories have investigated and documented current procedures for

monitoring and evaluation of post-tensioning to verify the effectiveness of inservice inspection 

examination methods for grouted and ungrouted tendon systems. Questions addressed include: 

(1) are current inservice inspection and examination methods adequate and (2) if corrective 

measures are used, are they adequate. In general, nondestructive evaluation and sensing 

methods are the only way to monitor the structural health of grouted tendons, and 

improvements to available methods should be investigated. Similarly, remediation of problems 

such as corrosion with grouted systems is difficult. The evaluation and repair of ungrouted 

prestressing systems is more straightforward.

v

FOREWORD

The prestressing systems used in containment vessels for commercial nuclear power plants are 

a critical structural element and must be protected from and monitored for any unanticipated 

degradation. Understanding the differences between grouted and ungrouted prestressing 

systems and the structural implications for nuclear containments is important for accurate 

containment analysis and for licensing. Corrosion of the prestressing tendons and anchorages 

represent a significant degradation mechanism for the serviceability of any prestressed concrete 

structure and for this reason various engineered approaches to tendon protection are employed

in the field. Injection of cementitious grout into the tendon ducts to encapsulate and protect the 

tendon is a common practice in the prestressing industry, yet there is limited experience with

servicing and maintaining grouted tendon systems for nuclear containment structures in the 

United States. Prestressed concrete containment structures in the United States use injected 

grease to protect the tendon systems from corrosion. 

The objective of this research is to investigate the structural behavior of a concrete containment 

vessel with grouted and ungrouted tendon systems. In addition, there is a need to document 

and compare inservice inspection requirements provided in peer reviewed standards and codes 

for grouted and ungrouted tendons. Finally, this report examines corrosion protection used for 

grouted and ungrouted tendons to assess the adequacy of the protection for the expected life of 

the structure.

vii

TABLE OF CONTENTS

Section Page

ABSTRACT............................................................................................................................... iii

FOREWORD ..............................................................................................................................v

TABLE OF CONTENTS ........................................................................................................... vii

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

LIST OF TABLES.................................................................................................................... xix

EXECUTIVE SUMMARY ......................................................................................................... xxi

ACKNOWLEDGMENTS.........................................................................................................xxiii

ACRONYMS AND ABBREVIATIONS.................................................................................... xxv

1 INTRODUCTION ..............................................................................................................1-1

1.1 Background ....................................................................................................... 1-1

1.2 Objective............................................................................................................ 1-2

1.3 The Approach .................................................................................................... 1-2

1.4 Overview of Industry Practice for Design and Severe Accident Analysis............ 1-3

2 COMPARISON OF STRUCTURAL RESPONSE OF GROUTED AND UNGROUTED 

TENDON SYSTEMS.........................................................................................................2-1

2.1 Comparison of Existing Methods of Modeling Grouted vs. Ungrouted Tendons. 2-2

2.1.1 Literature Review and Methods used by International Participants in 

Round-Robin Test Prediction Analysis Exercises................................. 2-2

2.1.2 Finite Element Program modeling Methods ......................................... 2-6

2.1.2.1 Basic Requirements ............................................................ 2-6

2.1.2.2 Specializations for Modeling PCCVs ................................... 2-6

2.1.2.3 Element Libraries ................................................................ 2-7

2.2 Recent (2010-2011) Developments in Tendon Modeling ................................... 2-8

2.2.1 Special Consideration of Anchorage Zones ......................................... 2-8

2.2.2 Comments on State-of-the-Practice for Modeling Tendons in Other 

Structures .......................................................................................... 2-12

2.2.3 Other Issues for Consideration – Reduced Bond of Tendons 

Compared to Reinforcement.............................................................. 2-15

2.3 Finite Element Models of PCCV with Ungrouted and Grouted Tendons........... 2-16

2.3.1 FE Representation of Rebar .............................................................. 2-16

2.3.2 FE Representation of Tendons .......................................................... 2-16

2.3.3 Simulating Initial Conditions and Losses............................................ 2-18

2.3.4 Material / Failure Criterion.................................................................. 2-29

2.4 Three Dimensional Global Finite Element Model ............................................. 2-31

2.4.1 Results and Comparisons from Prestressing Loads and Pressure 

Loads................................................................................................. 2-42

2.4.2 Grouted Tendon Model Method and Comparisons of Results............ 2-59

2.5 Conclusions ..................................................................................................... 2-67

3 COMPARISON OF POST-TENSIONING AND INSERVICE INSPECTION OF 

GROUTED AND UNGROUTED SYSTEM........................................................................3-1

3.1 Comparison of Post-Tensioning Systems .......................................................... 3-3

3.1.1 Post-Tensioning Methodology ............................................................. 3-3

3.1.2 Key Difference between Post-Tensioning Grouted and Ungrouted 

Tendon Systems.................................................................................. 3-8
viii

3.1.3 Industry and Structural Code Comparison of Post-Tensioning 

Methodologies ................................................................................... 3-10

3.1.4 Post-Tensioning Duct and Grouting Practice ..................................... 3-11

3.1.4.1 ACI 318 Requirements ...................................................... 3-11

3.1.4.2 Caltrans Requirements...................................................... 3-13

3.1.4.3 AASHTO Requirements .................................................... 3-14

3.1.4.4 FHWA Requirements ........................................................ 3-15

3.1.4.5 ASME Requirements......................................................... 3-21

3.1.4.6 US RG 1.107 Requirements for Grouted Tendons ............ 3-24

3.1.4.7 Comparison Table............................................................. 3-25

3.1.5 Summary and Conclusions ................................................................ 3-26

3.2 Comparison of Inservice Inspection Methods for Grouted and Ungrouted 

Tendon Systems.............................................................................................. 3-26

3.2.1 Needs of Post-Tensioned Monitoring................................................. 3-26

3.2.1.1 Evidence Provided by Bridge Industry on Need for 

Monitoring ......................................................................... 3-26

3.2.1.2 Evidence Provided by PCCV Industry on Need for 

Monitoring ......................................................................... 3-29

3.2.2 Different Industry and Code Requirements for Monitoring Loss of 

Prestress ........................................................................................... 3-31

3.2.2.1 Building and Waterfront Structure Industry Practice .......... 3-31

3.2.2.2 Nuclear Structure (PCCV) Practice ................................... 3-32

3.2.2.3 Regulatory Practice in Various Countries .......................... 3-35

3.2.2.4 Canada ............................................................................. 3-46

3.2.2.5 China................................................................................. 3-46

3.2.2.6 Czech Republic ................................................................. 3-47

3.2.2.7 Finland .............................................................................. 3-47

3.2.2.8 France............................................................................... 3-47

3.2.2.9 Japan ................................................................................ 3-47

3.2.2.10 Korea ................................................................................ 3-48

3.2.2.11 Sweden ............................................................................. 3-48

3.2.2.12 United Kingdom................................................................. 3-48

3.2.2.13 Summary of Comparison of French, Canadian, and United 

States Regulatory Practice Concerning Grouted Tendons. 3-48

3.2.3 PCCV Inservice Inspection Methods Related to Post-Tensioning ...... 3-49

3.2.3.1 Visual Inspection ............................................................... 3-50

3.2.3.2 Acoustic Monitoring........................................................... 3-50

3.2.3.3 Ultrasonic Pulse Echo ....................................................... 3-51

3.2.3.4 Fiber Optic Sensors........................................................... 3-53

3.2.3.5 Magnetic Methods............................................................. 3-54

3.2.3.6 Eddy Current..................................................................... 3-55

3.2.3.7 Ground Penetrating Radar ................................................ 3-55

3.2.3.8 Impact Echo ...................................................................... 3-56

3.2.3.9 Radiography...................................................................... 3-57

3.2.3.10 Vibrating Wire Strain Gauges ............................................ 3-57

3.2.4 Limitations of Monitoring for Grouted and Ungrouted Tendon 

Systems in PCCVs ............................................................................ 3-58

3.2.5 Corrective Measures in PCCVs ......................................................... 3-59

3.2.6 Summary and Conclusions ................................................................ 3-60

4 ASSESSMENT OF DURABILITY AND LONG-TERM CORROSION PROTECTION .......4-1

4.1 Introduction........................................................................................................ 4-1

4.2 Chemistry of Corrosion Process ........................................................................ 4-2
ix

4.3 Corrosion Over Time ......................................................................................... 4-5

4.3.1 Aging Mechanisms Related to Corrosion ............................................. 4-5

4.3.2 Comparison of General Lifetime of Tendon Systems as a Function of 

Corrosion............................................................................................. 4-6

4.3.3 Progression Characteristics of Corrosion in Each Type of System....... 4-7

4.4 Description of Long-Term Corrosion Protection Methods................................... 4-9

4.5 How Corrosion Affects Tendon Element Strength for Different Systems and 

Protection Methods.......................................................................................... 4-11

4.6 Tendon Sheathing Filler Characteristics and Effect on Corrosion Behavior...... 4-12

4.6.1 Cementitious Grout Characteristics ................................................... 4-12

4.6.2 Grease Characteristics ...................................................................... 4-12

4.7 Tendon Duct Type and Effect on Corrosion Behavior ...................................... 4-13

4.8 Stress State as a Function of Corrosion........................................................... 4-14

4.8.1 Loading.............................................................................................. 4-16

4.8.1.1 Results .............................................................................. 4-16

4.8.2 Conclusions....................................................................................... 4-34

4.9 Other Physical Conditions of PCCVs Affecting Corrosion and Durability.......... 4-38

4.10 FEA Modeling Tendon Corrosion for PCCVs and Predicted Effects on 

Performance.................................................................................................... 4-40

4.10.1 FEA Modeling .................................................................................... 4-40

4.10.2 Failure Criteria................................................................................... 4-45

4.10.3 FEA Simulation Results ..................................................................... 4-46

4.10.3.1 Corrosion Case 1 (Vertical Tendons near Wall-Base 

Juncture), Ungrouted......................................................... 4-46

4.10.3.2 Corrosion Case 1 (Vertical Tendons near Wall-Base 

Juncture), Grouted ............................................................ 4-51

4.10.3.3 Corrosion Case 1 (Vertical Tendons near Wall-Base 

Juncture) Summary........................................................... 4-54

4.10.3.4 Corrosion Case 2 (Hoop Tendons near E/H), Ungrouted... 4-56

4.10.3.5 Corrosion Case 2 (Hoop Tendons near E/H), Grouted ...... 4-66

4.10.3.6 Corrosion Case 2 (Hoop Tendons near E/H) Summary..... 4-70

4.10.3.7 Corrosion Case 1 and 2 Combined, Ungrouted................. 4-73

4.10.3.8 Corrosion Case 1 and 2 Combined, Grouted..................... 4-77

4.10.3.9 Corrosion Case 1 and 2 Summary .................................... 4-82

4.10.3.10 Corrosion Case 3 (At Anchor Zone), Ungrouted ................ 4-84

4.10.3.11 Corrosion Case 3 (At Anchor Zone), Grouted.................... 4-87

4.10.3.12 Corrosion Case 3 (At Anchor Zone) Summary .................. 4-91

4.10.3.13 Summary of Corrosion Cases 1, 2, and 3.......................... 4-93

4.10.4 FEA Simulation Results - Corrosion Applied After Pressure .............. 4-98

4.10.4.1 Corrosion Case 2 Applied at 1.5 and 2.5xPd, Ungrouted 

and Grouted ...................................................................... 4-98

4.10.4.2 Corrosion Case 3 Applied at 1.0, 1.5, and 2.5xPd, 

Ungrouted and Grouted................................................... 4-102

4.10.4.3 Summary of Corrosion Applied After Pressure Cases 2, 

and 3............................................................................... 4-110

4.11 Transitioning From FEA to Probabilistic Study for Corrosion.......................... 4-112

4.11.1 Pressure at Which Cylinder Stress Overcomes Prestress, Po ......... 4-113

4.11.2 Cylinder Hoop Cracking Pressure, Phc............................................ 4-114

4.11.3 Pressure at Liner Yield, Ply.............................................................. 4-115

4.11.4 Summary of Randomness and Uncertainties................................... 4-118

4.12 Detecting Corrosion in Grouted vs Ungrouted Systems ................................. 4-120

4.12.1 Inspection and Monitoring Techniques............................................. 4-121

4.12.2 Engineering Approach to Tendon Inspections.................................. 4-123
x

4.12.3 New Developments in Monitoring .................................................... 4-125

4.12.4 Repair of Tendons with Defective Grouting...................................... 4-126

4.12.5 Grouting of Voids............................................................................. 4-129

4.12.6 Closing of the Tendon...................................................................... 4-130

4.13 Summary Grouted and Ungrouted Tendons .................................................. 4-133

5 SUMMARY AND CONCLUSIONS....................................................................................5-1

5.1 Comparison of Structural Response of Grouted and Ungrouted Tendon 

Systems............................................................................................................. 5-1

5.2 Grouted and Ungrouted Tendon System Comparison ....................................... 5-2

5.3 Tendon Corrosion for Grouted and Ungrouted Systems .................................... 5-3

5.4 Overall Conclusions........................................................................................... 5-4

6 REFERENCES .................................................................................................................6-1
xi

LIST OF FIGURES

Figure Page

2-1 Axisymmetric model of 1:4-scale PCCV .......................................................................2-4

2-2 Local model of equipment hatch, 1:4-scale PCCV........................................................2-5

2-3 Three-dimensional cylinder mid-height model (3DCM), 1:4-scale PCCV ......................2-5

2-4 Schematic of 1:10 scale Sizewell model [SAND-90-2237C 1991].................................2-9

2-5 Section view of posttest condition of the 1:10-scale model [NUREG/CR-5671 

1998] .........................................................................................................................2-10

2-6 Comparison of pretest analyses vs. experiment results for 1:10 scale P/C model

[SAND-90-2237C 1991]. a) Radial displacements at midheight, b) Basemat uplift 

displacement, c) Hoop tendon strain at midheight, d) Meridional rebar strains. ..........2-11

2-7 Study of tendon pull-out problem in curved concrete post-tensioned box girders 

[NCHRP 2009]............................................................................................................2-13

2-8 Study of tendon anchorage zone cracking, view of typical FE mesh [NCHRP 2009]...2-14

2-9 Study of tendon anchorage zone cracking, view of tendon mesh [NCHRP 2009] .......2-15

2-10 Tendon ducts in 1:4-scale PCCV (silver colored tubes) along with traditional rebar....2-17

2-11 Tendon ducts (silver colored tubes) in 1:4-scale PCCV with and without 

connections as well as traditional rebar ......................................................................2-18

2-12 Method of simulating tendon friction using friction truss-ties .......................................2-21

2-13 H53 tendon force comparisons to pretest (From NUPEC/NRC PCCV test at SNL) ....2-22

2-14 Model-1 Abaqus Model...............................................................................................2-22

2-15 Tendon layout.............................................................................................................2-23

2-16 Anchorage of tendon to concrete................................................................................2-24

2-17 Tendon stress.............................................................................................................2-24

2-18 Tendon strain .............................................................................................................2-25

2-19 Tendon strain with (a) contact surfaces and (b) connector slots used to define 

contact........................................................................................................................2-26

2-20 Deformed shape with (a) contact surfaces and (b) connector slots used to define 

contact........................................................................................................................2-27

2-21 Tendon slip vs. location resulting from jacking the tendon from the un-stressed 

state (ungrouted)........................................................................................................2-28

2-22 Tendon slip vs. location after anchorage, during pressurization (ungrouted) ..............2-29

2-23 Tendon stress strain curve .........................................................................................2-31

2-24 The geometry for model 3 is based on the 1:4 scale PCCV test ( Figure 1.2 from 

[NUREG/CR-6810 2003]) ...........................................................................................2-31

2-25 Model 3 overview .......................................................................................................2-32

2-26 Meshed concrete vessel with various section assignments. Variations due to rebar 

layers and concrete thickness. Thickness of shell element rendered in ABAQUS ......2-33

2-27 Meshed concrete vessel with various section assignments. View of M/S and F/W .....2-33

2-28 Rigid links from bottom of vessel to basemat elements ..............................................2-34

2-29 Vertical tendon jacking element ends rigid linked to closest basemat node ................2-34

2-30 Hoop tendon jacking element ends tied to closest buttress center node.....................2-35

2-31 Hoop tendon jacking elements (‘nubs’) and tendon nodes shown relative to 

concrete nodes in the buttress region.........................................................................2-36

2-32 Deformed shape of hoop tendon anchor system after “jacking loading step”..............2-37

2-33 Deformed shape of hoop tendon system after “anchor set step”.................................2-38

2-34 Hoop tendon layout ....................................................................................................2-39

2-35 Vertical tendon layout.................................................................................................2-40

2-36 Calculated tendon stress profile with losses for two tendons at given azimuth (blue 

and purple line)...........................................................................................................2-41
xii

2-37 Deformed shape after tendon anchorage (deformation scale × 500) ..........................2-43

2-38 Deformed shape at 3.6 × Pd (deformation scale × 20) ................................................2-44

2-39 Deformed shape at anchoring at elev. 4.68 m (15’-4 1/16”) (x500) .............................2-45

2-40 Deformed shape at Pd at elev. 4.68 m (15’-4 1/16”) (x500).........................................2-45

2-41 Deformed shape at 1.5 x Pd at elev. 4.68 m (15’-4 1/16”) (x250) ...............................2-46

2-42 Deformed shape at 2.0 x Pd at elev. 4.68 m (15’-4 1/16”) (x100) ................................2-46

2-43 Deformed shape at 2.5 x Pd at elev. 4.68 m (15’-4 1/16”) (x50) ..................................2-47

2-44 Deformed shape at 3.0 x Pd at elev. 4.68 m (15’-4 1/16”) (x30) ..................................2-47

2-45 Deformed shape at 3.4 x Pd at elev. 4.68 m (15’-4 1/16”) (x30) ..................................2-48

2-46 Deformed shape at 3.6 x Pd at elev. 4.68 m (15’-4 1/16”) (x30) ..................................2-48

2-47 Maximum principal membrane strain in concrete at 3.6 x Pd ......................................2-49

2-48 Stress in hoop tendons anchored at 90° after jacking before anchorage ....................2-50

2-49 Stress in hoop tendons anchored at 90° after anchorage ...........................................2-51

2-50 Stress in vertical tendons after jacking before anchorage (contours in psi).................2-52

2-51 Stress in vertical tendons after anchorage (contours in psi)........................................2-53

2-52 Abaqus analysis – hoop tendon H35 force ................................................................2-54

2-53 LST test – hoop tendon H35 force..............................................................................2-54

2-54 SFMT test – hoop tendon H35 force...........................................................................2-55

2-55 Abaqus analysis – hoop tendon H53 force .................................................................2-55

2-56 LST test – hoop tendon H53 force..............................................................................2-56

2-57 SFMT test – hoop tendon H53 force...........................................................................2-56

2-58 Abaqus analysis – hoop tendon H68 force .................................................................2-57

2-59 LST test – hoop tendon H68 force..............................................................................2-57

2-60 SFMT test – hoop tendon H68 force...........................................................................2-58

2-61 Comparison of model 3 (ungrouted) result to LST test measurement at standard 

output location 14 which is the radial displacement at the center of the equipment 

hatch. Pressure is MPa with the grid divisions representing multiples of the design 

pressure (0.39 MPa)...................................................................................................2-59

2-62 During jacking – slot connector element connecting tendons to vessel.......................2-60

2-63 After anchorage– slot connector element removed and replaced with beam 

connector elements ....................................................................................................2-60

2-64 Comparison of grouted vs. ungrouted tendon model 3 results for radial 

displacements at elev. 4.68 meters (near cylinder mid-height) ...................................2-61

2-65 Comparison of grouted vs. ungrouted tendon model 3 results for tendon H35 

strains.........................................................................................................................2-62

2-66 Comparison of grouted vs. ungrouted tendon model 3 results for tendon H53 

strains.........................................................................................................................2-63

2-67 Comparison of grouted vs. ungrouted tendon model 3 results for tendon H68 

strains.........................................................................................................................2-64

2-68 Comparison of (a) grouted and (b) ungrouted results from model 3 for tendon V37, 

Azimuth 240 o .............................................................................................................2-65

2-69 Comparison of (a) grouted and (b) ungrouted results from model 3 for tendon V46, 

Azimuth 135 o .............................................................................................................2-66

3-1 Typical ungrouted single strand tendon ........................................................................3-4

3-2 Typical post-tensioning detail for bridges or other civil structures .................................3-5

3-3 Typical post-tensioning detail for PCCV from [NUREG/CR-6810 2003] for (a) 

Vertical section (b) Horizontal section...........................................................................3-6

3-4 PCCV wall mock-up from [NUREG/CR-6810 2003]......................................................3-7

3-5 Summary of potential tendon relaxation demonstrating the decrease in relaxation 

time for increasing applied stress [Structural Research Series No. 237 1962]. ...........3-10

3-6 Spiral Wound Steel Duct and Rigid Steel Pipe............................................................3-17

3-7 Corrugated plastic duct...............................................................................................3-17

3-8 Corrosion observed near ends of tendons from the Sunshine Skyway Bridge ............3-28
xiii

3-9 Energy vs. time of wire breaks in 1:4 scale SMFT [NUREG/CR-6810 2003]...............3-51

3-10 Ultrasonic inspection setup [Lafhaj, Goueygou et al. 2006].........................................3-52

3-11 Basic structure of a fiber optic sensor.........................................................................3-53

3-12 Depiction of a Vibrating Wire Strain Gauge ................................................................3-58

4-1 Corrosion of reinforcement – an electrochemical process [Collins and Mitchell 

1997] ...........................................................................................................................4-3

4-2 Model 1 with Corroded Tendon at Anchorage.............................................................4-15

4-3 Reduction in Effective Area due to Corrosion .............................................................4-15

4-4 Strain in corroded tendon for ungrouted case @ 1.5xPd ............................................4-17

4-5 Strain in uncorroded tendon (opposite) for ungrouted case @ 1.5xPd........................4-17

4-6 Strain in liner for ungrouted case @ 1.5xPd ...............................................................4-18

4-7 Displacement of liner for ungrouted case @ 1.5xPd ...................................................4-18

4-8 Strain in corroded tendon for grouted case @ 1.5xPd ................................................4-19

4-9 Strain in uncorroded tendon for grouted case @ 1.5xPd ............................................4-20

4-10 Strain in liner for grouted case @ 1.5xPd ...................................................................4-20

4-11 Displacement of liner for grouted case @ 1.5xPd .......................................................4-21

4-12 Strain in corroded tendon for ungrouted case @ 1.0xPd ............................................4-22

4-13 Strain in uncorroded tendon for ungrouted case @ 1.0xPd.........................................4-22

4-14 Strain in liner for ungrouted case @ 1.0xPd ...............................................................4-23

4-15 Displacement of liner for ungrouted case @ 1.0xPd ...................................................4-23

4-16 Strain in uncorroded tendon for grouted case @ 1.0xPd ............................................4-24

4-17 Strain in corroded tendon for grouted case @ 1.0xPd ................................................4-24

4-18 Strain in liner for grouted case @ 1.0xPd ...................................................................4-25

4-19 Displacement of liner for grouted case @ 1.0xPd .......................................................4-25

4-20 Strain in corroded tendon for ungrouted case @ 2.5xPd ............................................4-26

4-21 Strain in uncorroded tendon for ungrouted case @ 2.5xPd.........................................4-26

4-22 Strain in liner for ungrouted case @ 2.5xPd ...............................................................4-27

4-23 Displacement of liner for ungrouted case @ 2.5xPd ...................................................4-27

4-24 Strain in corroded tendon for grouted case @ 2.5xPd ................................................4-28

4-25 Strain in uncorroded tendon for grouted case @ 2.5xPd ............................................4-28

4-26 Strain in liner for grouted case @ 2.5xPd ...................................................................4-29

4-27 Displacement of liner for grouted case @ 2.5xPd .......................................................4-29

4-28 Strain in corroded tendon for ungrouted case @ 3.0xPd ............................................4-30

4-29 Strain in liner for ungrouted case @ 3.0xPd ...............................................................4-30

4-30 Strain in liner for ungrouted case @ 3.0xPd ...............................................................4-31

4-31 Displacement of liner for ungrouted case @ 3.0xPd ...................................................4-31

4-32 Strain in corroded tendon for grouted case @ 3.0xPd ................................................4-32

4-33 Strain in liner for ungrouted case @ 3.0xPd ...............................................................4-32

4-34 Strain in liner for grouted case @ 3.0xPd ...................................................................4-33

4-35 Displacement of liner for grouted case @ 3.0xPd .......................................................4-33

4-36 Strain in the corroded tendon adjacent to the corroded section which demonstrates 

the unloading of the tendon as the corroded section elongates ..................................4-34

4-37 Maximum Strain in Anywhere in Uncorroded Tendon which demonstrates the load 

transfer from the elongating corroding tendon to the uncorroded tendon....................4-35

4-38 Maximum Liner Strain which demonstrates the degree to which the containment 

deforms in response to the corroding tendon..............................................................4-36

4-39 Radial Displacement of Liner which shows the response of the containment to the 

loss of support from the corroding tendon...................................................................4-36

4-40 Postulated corrosion Zones ........................................................................................4-42

4-41 Corrosion Case 1 - Highlighted tendons indicate corroded areas along the bottom 

of the wall ...................................................................................................................4-43

4-42 Corrosion Case 2 - Highlighted tendons indicate corroded areas near the 

equipment hatch.........................................................................................................4-44
xiv

4-43 Corrosion Case 3 - Highlighted tendons indicate corroded areas at anchor zone .......4-44

4-44 Stress-Strain Properties for Tendon ...........................................................................4-45

4-45 Stress-Strain Properties for Liner................................................................................4-45

4-46 Stress in Vertical Tendons A) after Anchorage, and B) after Corrosion Case 1 

(Note Contours in units of psi) ....................................................................................4-46

4-47 Stress in Vertical Tendons in Dome A) after Anchorage, and B) after Corrosion 

Case 1 (Note Contours in units of psi) ........................................................................4-47

4-48 Maximum Principal Strain in Liner at 1.5 x Pd, Corrosion Case 1, Ungrouted.............4-48

4-49 Maximum Principal Strain in Liner at 3.3 x Pd, Corrosion Case 1, Ungrouted 

(Contours Modified)....................................................................................................4-48

4-50 Radial displacement comparisons between uncorroded, ungrouted and corrosion 

case 1 (Vertical Tendons at basemat-wall juncture, approximately 0° azimuth), 

ungrouted...................................................................................................................4-49

4-51 Locations for strains over selected gage length near E/H...........................................4-50

4-52 Strains over selected gage length near E/H, corrosion case 1, ungrouted..................4-51

4-53 Maximum Principal Strain in Liner at 1.5 x Pd, Corrosion Case 1, Grouted ................4-52

4-54 Maximum Principal Strain in Liner at 3.6 x Pd, Corrosion Case 1, Grouted ................4-53

4-55 Radial displacement comparisons between uncorroded, grouted and corrosion 

case 1 (Vertical Tendons at basemat-wall juncture, approximately 0° azimuth), 

grouted.......................................................................................................................4-53

4-56 Strains over Selected Gage Length Near E/H, Corrosion Case 1, Grouted ................4-54

4-57 Maximum Liner Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted..............................................................................................4-55

4-58 Maximum tendon strain comparisons between uncorroded and corroded, 

ungrouted and grouted ...............................................................................................4-56

4-59 Stress in hoop tendons anchored at 90° after anchorage, corrosion case 2 (Note

units of psi).................................................................................................................4-57

4-60 Stress in hoop tendons anchored at 270° after anchorage, corrosion case 2 (Note 

units of psi).................................................................................................................4-58

4-61 Stress in hoop tendons anchored at 90° after anchorage and corrosion, corrosion 

case 2 (Note units of psi)............................................................................................4-58

4-62 Stress in Hoop Tendons Anchored at 270° after Anchorage and Corrosion, 

Corrosion Case 2 (Note units of psi)...........................................................................4-59

4-63 Strain in hoop tendons anchored at 90° after anchorage, corrosion case 2 ................4-59

4-64 Strain in hoop tendons anchored at 270° after anchorage, corrosion case 2 ..............4-60

4-65 Strain in Hoop Tendons Anchored at 90° after Anchorage and Corrosion, 

Corrosion Case 2 .......................................................................................................4-60

4-66 Strain in hoop tendons anchored at 270° after anchorage and corrosion, corrosion 

case 2.........................................................................................................................4-61

4-67 Strain in hoop tendons anchored at 90° at 2.8 x Pd, corrosion case 2, ungrouted ......4-61

4-68 Strain in hoop tendons anchored at 270° at 2.8 x Pd, corrosion case 2, ungrouted ....4-62

4-69 Maximum principal strain in liner at 1.5 x Pd, corrosion case 2, ungrouted.................4-62

4-70 Maximum principal strain in liner at 2.8 x Pd, corrosion case 2, ungrouted.................4-63

4-71 Maximum principal strain in liner at 2.8 x Pd, showing deformed shape x30, 

corrosion case 2, ungrouted .......................................................................................4-64

4-72 Radial Displacement Comparisons between Uncorroded, Ungrouted and 

Corrosion Case 2 (Hoop tendons near the equipment hatch), Ungrouted...................4-65

4-73 Liner Strains over Selected Gage Length Near E/H, Corrosion Case 2. Ungrouted....4-66

4-74 Strain in hoop tendons anchored at 90° at 3.6 x Pd, corrosion case 2, grouted ..........4-67

4-75 Strain in hoop tendons anchored at 270° at 3.6 x Pd, corrosion case 2, grouted ........4-68

4-76 Maximum principal strain in liner at 1.5 x Pd, corrosion case 2, grouted.....................4-68

4-77 Maximum principal strain in liner at 3.6 x Pd, corrosion case 2, grouted.....................4-69
xv

4-78 Radial displacement comparisons between uncorroded, grouted and corrosion 

case 2 (Hoop tendons near the equipment hatch), grouted ........................................4-69

4-79 Local Liner Strains over selected gage length near E/H, corrosion case 2, grouted....4-70

4-80 Maximum Radial Displacement Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted..............................................................................................4-71

4-81 Maximum Liner Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted..............................................................................................4-72

4-82 Maximum Tendon Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted..............................................................................................4-72

4-83 Strain in Hoop Tendon Anchored at 90° at 2.3xPd, Corrosion Case 1&2, 

Ungrouted (Final Step) ...............................................................................................4-73

4-84 Strain in Hoop Tendon Anchored at 270° at 2.3xPd, Corrosion Case 1&2, 

Ungrouted(Final Step) ................................................................................................4-74

4-85 Maximum Principal Strain in Liner at 1.0xPd, Corrosion Case 1&2, Ungrouted ..........4-74

4-86 Maximum Principal Strain in Liner at 2.3xPd, Corrosion Case 1&2, Ungrouted 

(Final Step).................................................................................................................4-75

4-87 Radial displacement comparisons between uncorroded, ungrouted and corrosion 

case 1 and 2 (Vertical tendons at basemat-wall juncture and hoop tendons near 

the equipment hatch), ungrouted................................................................................4-76

4-88 Strains over Selected Gage Length Near E/H, Corrosion Case 1 and 2, UnGrouted ..4-77

4-89 Strain in Hoop Tendons Anchored at 90° at 2.3xPd, Corrosion Case 1&2, Grouted ...4-78

4-90 Strain in Hoop Tendons Anchored at 270° at 2.3xPd, Corrosion Case 1&2, 

Grouted ......................................................................................................................4-78

4-91 Strain in hoop tendons anchored at 90° at 3.6xPd, corrosion case 1&2, grouted........4-79

4-92 Strain in hoop tendons anchored at 270° at 3.6xPd, corrosion case 1&2, grouted......4-79

4-93 Maximum principal strain in liner at 1.0xPd, corrosion case 1&2, grouted...................4-80

4-94 Maximum principal strain in liner at 2.3xPd, corrosion case 1&2, grouted...................4-80

4-95 Maximum principal strain in liner at 3.6xPd, corrosion case 1&2, grouted...................4-81

4-96 Radial displacement comparisons between uncorroded, grouted and corrosion 

case 1 and 2 (Vertical tendons at basemat-wall juncture and hoop tendons near 

the equipment hatch), grouted....................................................................................4-81

4-97 Local Liner Strains over Selected Gage Length Near E/H, Corrosion Case 1 and 2, 

Grouted ......................................................................................................................4-82

4-98 Maximum radial displacement comparisons between uncorroded and corroded, 

ungrouted and grouted ...............................................................................................4-83

4-99 Maximum liner strain comparisons between uncorroded and corroded, ungrouted 

and grouted ................................................................................................................4-83

4-100 Maximum tendon strain comparisons between uncorroded and corroded, 

ungrouted and grouted ...............................................................................................4-84

4-101 Strain in Hoop Tendons Anchored at 90° at 3.2xPd, Corrosion Case 3, Ungrouted 

(Final Step).................................................................................................................4-85

4-102 Strain in Hoop Tendons Anchored at 270° at 3.2xPd, Corrosion Case 3, Ungrouted 

(Final Step).................................................................................................................4-85

4-103 Maximum Principal Strain in Liner at 1.0xPd, Corrosion Case 3, Ungrouted...............4-86

4-104 Maximum principal strain in liner at 3.2xPd, corrosion case 3, ungrouted (final 

step) .........................................................................................................................4-86

4-105 Radial displacement comparisons between uncorroded, ungrouted and corrosion 

case 3 (Horizontal Tendons near anchorage, approximatel 260° azimuth), 

ungrouted...................................................................................................................4-87

4-106 Strain in hoop tendons anchored at 90° at 3.2xPd, corrosion case 3, grouted ............4-88

4-107 Strain in hoop tendons anchored at 270° at 3.2xPd, corrosion case 3, grouted ..........4-88

4-108 Strain in hoop tendons anchored at 90° at 3.6xPd, corrosion case 3, grouted ............4-89

4-109 Strain in hoop tendons anchored at 270° at 3.6xPd, corrosion case 3, grouted ..........4-89
xvi

4-110 Maximum principal strain in liner at 1.0xPd, corrosion case 3, grouted.......................4-90

4-111 Maximum principal strain in liner at 3.6xPd, corrosion case 3, grouted.......................4-90

4-112 Radial displacement comparisons between uncorroded, grouted and corrosion 

case 3 (Horizontal Tendons near anchorage, approximately 260° azimuth), 

grouted.......................................................................................................................4-91

4-113 Maximum radial displacement comparisons between uncorroded and corroded, 

ungrouted and grouted ...............................................................................................4-92

4-114 Maximum liner strain comparisons between uncorroded and corroded, ungrouted 

and grouted ................................................................................................................4-92

4-115 Maximum tendon strain comparisons between uncorroded and corroded, 

ungrouted and grouted ...............................................................................................4-93

4-116 Comparison of Radial Displ. at E/H – All Corrosion Cases, Grouted & Ungrouted 

vs. Uncorroded...........................................................................................................4-94

4-117 Radial displacement comparisons between uncorroded, ungrouted and corrosion 

at 1.5Pd Case 2 (hoop tendons near the equipment hatch).........................................4-99

4-118 Radial displacement comparisons between uncorroded, ungrouted and corrosion 

at 2.5Pd Case 2 ..........................................................................................................4-99

4-119 Radial displacement comparisons between uncorroded, grouted and corrosion at 

1.5Pd Case 2 ............................................................................................................4-100

4-120 Radial displacement comparisons between uncorroded, grouted and corrosion at 

2.5Pd Case 2 ............................................................................................................4-100

4-121 Maximum Radial Displacement Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted............................................................................................4-101

4-122 Maximum Liner Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted............................................................................................4-101

4-123 Maximum Tendon Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted............................................................................................4-102

4-124 Radial Displacement Comparisons between Uncorroded, Ungrouted and 

Corrosion at 1.0xPd Case 3 ......................................................................................4-103

4-125 Radial Displacement Comparisons between Uncorroded, UnGrouted and 

Corrosion at 1.5xPd Case 3, Ungrouted....................................................................4-103

4-126 Radial Displacement Comparisons between Uncorroded, UnGrouted and 

Corrosion at 2.5xPd Case 3, Ungrouted....................................................................4-104

4-127 Radial Displacement Comparisons between Uncorroded, Grouted and Corrosion 

at 1.0xPd Case 3, Grouted........................................................................................4-104

4-128 Radial Displacement Comparisons between Uncorroded, Grouted and Corrosion 

at 1.5xPd Case 3, Grouted........................................................................................4-105

4-129 Radial Displacement Comparisons between Uncorroded, Grouted and Corrosion 

at 2.5xPd Case 3, Grouted........................................................................................4-105

4-130 Radial Displacement Comparisons between Uncorroded, Ungrouted and 

Corrosion at 2.5xPd Case 3, Ungrouted (Scale Magnified) .......................................4-106

4-131 Radial Displacement Comparisons between Uncorroded, Grouted and Corrosion 

at 2.5xPd Case 3, Grouted (Scale Magnified) ...........................................................4-107

4-132 Maximum Radial Displacement Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted............................................................................................4-108

4-133 Average Radial Displacement Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted............................................................................................4-108

4-134 Maximum Liner Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted............................................................................................4-109

4-135 Maximum Tendon Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted............................................................................................4-109

4-136 Approximate strain versus pressure response for cylinder hoop tendons in a 

PCCV .......................................................................................................................4-114
xvii

4-137 Tendon stress-strain curve........................................................................................4-115

4-138 Demonstration of approximate method for calculating global hoop tendon strain 

and liner strain..........................................................................................................4-119

4-139 Sample global strain curve with confidence intervals ................................................4-120

4-140 Demonstration of application of hypothetical corrosion to hoop tendons, and the 

influence on global liner strain ..................................................................................4-120

4-141 Typical inspection Strategy for a Box Girder Bridge..................................................4-124

4-142 Gaining Access to a Tendon [VSL Report Series 5 2002].........................................4-128

4-143 Vacuum Control and Grouting Equipment for Tendons with Defective Grouting .......4-130

4-144 Closing of Tendon Openings [VSL Report Series 5 2002] ........................................4-131

xix

LIST OF TABLES

Table Page

2-1 Summary of analytical models used for prior PCCV analysis studies [EPRI 

NP6263-SD 1989, NUREG/CR-6433 1996, NUREG/CR-5671 1998, NUREG/CR6685 2000, NUREG/CR-6809 2003].............................................................................2-3

2-2 Tendon stress distribution for standard tendon behavior analysis (includes seating 

losses and assumed linearly varying with azimuth in between points.).......................2-41

3-1 Physical properties required for shrink sleeves...........................................................3-19

3-2 Comparison of key parameters for ducts and grouting................................................3-25

3-3 Regulatory agencies and codes/guidance. .................................................................3-36

3-4 Reactors, containments, and pre-stressing systems used around the world...............3-39

3-5 Inspection methods and acceptance criteria for various countries. .............................3-41

3-6 Frequency of inspection and codes related to testing for various countries. ...............3-45

3-7 Summary of NDT methods and their applicability to monitoring prestressed 

PCCVs. ......................................................................................................................3-59

4-1 Allowable concrete crack widths from [CEB-FIP 1978]. ................................................4-5

4-2 Maximum liner strain (%) at 2.5xPd ............................................................................4-37

4-3 Affected azimuth at 2.5xPd (degrees) .........................................................................4-37

4-4 List of Model 3 Analyses.............................................................................................4-43

4-5 Maximum Tendon Strain for Uncorroded Cases at Pressure Milestones ....................4-94

4-6 Maximum Tendon Strain for Corrosion Case 1 at Pressure Milestones ......................4-95

4-7 Maximum Tendon Strain for Corrosion Case 2 at Pressure Milestones ......................4-95

4-8 Maximum Tendon Strain for Corrosion Case 3 at Pressure Milestones ......................4-96

4-9 Maximum Tendon Strain for Combined Corrosion Cases 1&2 at Pressure 

Milestones ..................................................................................................................4-96

4-10 Maximum Tendon Strain for Corrosion Case 2 @ 1.5xPd ........................................4-110

4-11 Maximum Tendon Strain for Corrosion Case 2 @ 2.5xPd ........................................4-110

4-12 Maximum Tendon Strain for Corrosion Case 3 @ 1.0xPd ........................................4-110

4-13 Maximum Tendon Strain for Corrosion Case 3 @ 1.5xPd ........................................4-111

4-14 Maximum Tendon Strain for Corrosion Case 3 @ 2.5xPd ........................................4-111

xxi

EXECUTIVE SUMMARY

There is a need for improved understanding of the behavior of grouted tendons in prestressed 

concrete containment vessels (PCCVs). Prior to 2010, there has been only one operating plant 

with a containment vessel in the United States with grouted tendons; these grouted tendons are 

only in the vertical tendons of the structure. With the influx of new plant applications, there are 

multiple licensees requesting the use of grouted tendon systems in the PCCVs. As a result the 

USNRC needs to evaluate the adequacy of the proposed grouted systems and any potential 

impacts on the safety of these new plants. This report provides a basis that objective. 

This NUREG consists of three sections: (1) an investigation into the structural behavior of 

PCCVs with grouted and ungrouted tendon systems, (2) a comparison of post-tensioning and in 

service inspection methods for containment vessels with grouted and ungrouted tendons, and 

(3) an assessment of the durability and long-term corrosion protection for grouted and 

ungrouted tendons.

The focus of the first section is to study the structural behavior, strength, and expected failure 

modes of a post-tensioned containment with grouted and ungrouted tendon systems. This work 

was done in parallel with the Organization of Economic and Cooperative Development / Nuclear 

Energy Agency (OECD/NEA) “Study on Post-Tensioning Methodologies in Containments,” 

[NEA 2015] where it has been proposed that methods be compared in a round robin study of 

advantages and disadvantages of grouted tendons. The OECD/NEA study builds on the 

International Standard Problem 48, “Containment Integrity.” A novel technique for modeling 

grouted tendon systems for prestressed concrete containment vessels (PCCVs) was developed 

that captures the initial stress distribution in the tendons that results from jacking and anchoring. 

The developed approach is more analytically accurate and computationally efficient than 

previous attempts, yet shows good agreement with previously accepted approaches and with 

the available PCCV test data. The results of the modeling effort indicate somewhat lower 

performance for grouted tendon systems when the PCCV is subjected to internal 

overpressurization. This reduction in capacity is attributed to the inability of the tendons in the 

grouted system to slip and redistribute loading in response to the internal pressure. By 

preventing the slippage of the tendons, stress concentrations arising from jacking and anchoring 

as well as local discontinuities in the model remain concentrated, thus leading to failure due to 

tendon rupture. In the ungrouted simulations, tendon movement reduced these maximum stress 

areas somewhat which resulted in increased ultimate structural capacity.

The focus of the second section is the investigation and documentation of current procedures 

for post-tensioning and monitoring of post-tensioning to verify the effectiveness of the inservice

inspection examination methods for grouted and ungrouted tendon systems. Such questions 

addressed are (1) are current examination methods adequate and (2) if corrective measures are 

used, are these measures adequate? The assessment of available NDE and monitoring 

methods for grouted prestressing systems concluded that more technical development will be 

required to achieve the same level of investigation confidence that exists with ungrouted 

systems. Various other countries are exploring these NDE and sensing techniques presently 

and it is expected that the United States will benefit from other countries research in these 

areas. Corrective measures for ungrouted tendon systems are much more readily available than 

for grouted systems. Some destructive repair techniques are possible for localized tendon 

problems, but executing these repairs depends on the ability to locate the problem with a high 

degree of certainty. 

The focus of the third section is the assessment of long-term corrosion protection methods to 

ensure life-long durability for the grouted and ungrouted tendon systems used in current and 

future nuclear power or related industries. A review and discussion of information and / or data 
xxii

related to corrosion that concerns tensioning element strength and ductility, tendon sheathing 

filler characteristics and tendon duct type is provided. Other aspects (e.g., aging mechanisms, 

stress) of material physical condition were addressed to assess durability. The Task 3 work also 

includes detailed finite element analysis studies, using two kinds of 3-dimensional models: 

small models representing only two hoop tendons, and large models (“Model 3”) of the complete 

PCCV 1:4 Scale containment structure.

The full, global 3D models showed that for the postulated corrosion cases studied, grouted 

tendons provide a significant structural advantage over ungrouted tendons. This advantage is 

on the basis of ultimate structural capacity in response to internal overpressurization, not impact 

or seismic. When the tendon corrosion was applied, the adjacent regions of the containment 

structure experienced some loss of prestressing. This effect was much more localized in the 

case of the grouted tendon systems and can be credited for the increased ultimate structural 

capacity. Conversely, the ungrouted tendon model lost prestressing over a much larger area 

when tendon corrosion was applied and this more widespread loss of prestress support resulted 

in reduced ultimate structural capacity. More specifically for ungrouted tendon models, when the 

vertical tendons near the basemat-wall juncture were corroded, ultimate capacity was reduced 

by about 10%. When hoop tendons near the equipment hatch were corroded, a 25% reduction 

in capacity was observed. When hoop tendon corrosion was applied near the anchorages a 

10% reduction in ultimate capacity was predicted. For these three cases with grouted tendon 

systems, the ultimate capacity was predicted to be similar to the uncorroded, grouted state, 

though with larger maximum tendon strains. 

As would be expected from pressure vessel theory, the modeling efforts indicated that hoop 

tendon corrosion was much more important to ultimate capacity than vertical tendon corrosion. 

This difference is particularly pronounced when the corroded area corresponds to areas of the 

vessel that experience greater hoop stress, such as near the equipment hatch. Hoop tendon 

corrosion near the anchorages, which would be an expected location for moisture ingress, was 

not found to be extremely penalizing because the buttress area has high intrinsic stiffness and 

lower tendon force owing to losses associated with anchor set.
xxiii

ACKNOWLEDGMENTS

The authors would like to acknowledge the continuing support and guidance throughout the 

program provided by U.S. Nuclear Regulatory Commission Project Manager and Technical

Monitor, Madhumita Sircar. In addition, the authors would like to thank all the organizations that 

are participating in the OECD WGIAGE Expert meeting. Their contributions to furthering the 

current state of the art in grouted tendons make this effort possible. 

For the investigation of corrosion protection methods and the study of corrosion scenarios of 

grouted and ungrouted tendon systems, the authors wish to thank Jessica Salazar and Cori 

Wallace of Sandia, and Patrick Chang and Christina Thung of Moffatt & Nichol.

Thomas Herrity of NRC is recognized for his assistance in the preparation of the final 

manuscript.

xxv

ACRONYMS AND ABBREVIATIONS

3D Three Dimensional

AASHTO American Association of State Highway and Transportation Officials

A/L Air Lock

Ag Gross Area of Section

BWR Boiling Water Reactor

CALTRANS State of California Department of Transportation

CEGB Central Electricity Generating Board

DG Draft Regulatory Guide

E/H Equipment Hatch

Ec,s Young’s Modulus of Concrete or Steel Respectively

EDF Électricité de France

EPR European Pressurized Reactor

εs Tensile Strain

F/W Feed Water

FBR Fast Breeder Reactor

fc’ Compressive Strength

fcp Compressive Prestress

fcr Cracking Strength

FEA Finite Element Analysis

FHWA US Department of Transportation, Federal Highway Administration

fmember Average Member Stress at Member Cracking

fp Prestressing Stress

Fp Force of Prestressing

GCR Gas Cooled Reactor

LST Limit State Test

LWGR Light Water Graphite Reactor
xxvi

M/S Main Steam

M&N Moffatt and Nichol

MDEP Multinational Design Evaluation Program

NDE Non-Destructive Evaluation

NDT Non-Destructive Testing

NEA Nuclear Energy Agency

NRC Nuclear Regulatory Commission

NUPEC Nuclear Power Engineering Corporation

OECD Organization for Economic Co-operation and Development

P/C Prestressed Concrete

PCCV Prestressed Concrete Containment Vessel

Pd Design Pressure 

Pjack Tendon Forces Achieved During Stressing

PHWR Pressurized Heavy Water Reactor

PWR Pressurized Water Reactor

R/C Reinforced Concrete

RG Regulatory Guide

SFMT Structural Failure Mode Test

SNL Sandia National Laboratories

SPE Specific Problem Exercise

STUK Radiation and Nuclear Safety Authority (Finland)

Ts Tensile Strength Contribution from Steel

UKEPR United Kingdom European Pressurized Reactor

WGIAGE Working Group on Integrity and Aging of Components

ρp Post-Tensioning Percentage

ρs Reinforcing Ratio of Rebar
1-1

1 INTRODUCTION

The United States Nuclear Regulatory Commission (USNRC) is investigating use of grouted 

versus ungrouted tendon systems in prestressed concrete containment vessels. The NRC 

investigation has three areas of focus. Sandia National Laboratories (SNL), with the support of 

Robert Dameron from Moffatt & Nichol (M&N), is assisting in this investigation. The first area of 

interest is a comparison of the structural behavior of concrete containment vessels with grouted 

and ungrouted tendon systems. The second focus of the investigation is a comparison of posttensioning methods and inservice inspection requirements for grouted and ungrouted tendons. 

Finally, long-term corrosion protection methods to ensure life-long durability for grouted and 

ungrouted tendon systems were assessed.

1.1 Background

As of fall 2014, there are 100 current operating nuclear power plants in the United States. 

Seventy two of these have concrete containments, and 37 are post-tensioned (only one

operating plant has some grouted or bonded tendons). In the U.S., post-tensioned concrete 

containment vessels (PCCVs) constitute the single largest class of containment structures. 

Although only one operating plant (H.B. Robinson) in the U.S. uses grouted tendons (vertical 

tendons), grouted systems have been used extensively in other parts of the world (Belgium, 

Canada, China, France, and Korea). Since post-tensioned containments with grouted tendons 

have been used in other countries and applicants for new plants in the U.S. have proposed to 

use grouted tendons, a research program is underway, to compare grouted and ungrouted 

prestressing systems. There is a need to compare the structural response of the two different 

types of systems, as well as to assess methods for ensuring the long term health of the fleet 

using either type of system. This research continues from the research done to date at SNL for

the USNRC in the field of containment structural integrity.

Containment buildings for nuclear power plants have been the subject of past studies and 

research efforts owing to their significant safety role as the final barrier containing radionuclide 

release during an accident. While the contributions of each of these efforts to the understanding 

of the role of containment in ensuring the safe operation of nuclear power plants is important, 

the most comprehensive experimental effort has been conducted at Sandia National 

Laboratories (SNL), primarily under the sponsorship of the United States Nuclear Regulatory 

Commission (NRC). NUREG/CR-6906, “Containment Integrity Research at Sandia National 

Laboratories: An Overview,” summarizes the major results of the experimental efforts as well as

the observations and insights gained from the analytical efforts for more than 25 years of 

containment integrity research at SNL. Prior to pressure testing the scale models, a number of 

regulatory and research organizations were invited to participate in pre-test Round Robin 

analyses to perform predictive modeling of the response of scale models to over pressurization. 

Many domestic and international organizations responded and agreed to participate in the pretest Round Robin analysis activities. The purpose of the Containment Integrity Research at SNL 

was to provide a forum for researchers in the area to apply current state-of-the-art analysis 

methodologies to predicting capacity of steel, reinforced, and prestressed concrete containment 

vessels. 

As noted above, this work is related to the NRC-sponsored containment integrity programs at 

SNL. These programs investigated the behavior of light water reactor (LWR) containment 

buildings under loadings that exceed the design basis (commonly referred to as "severe" 

accident loading). A combination of experimental and analytical studies was employed in these 

programs. Initially, over-pressurization tests of several scale model containment buildings were 

conducted [NUREG/CR-4216 1986, SAND--84-2153 1992, NUREG/CR-5121 1992, SAND--98-
1-2

1044C 1998, NUREG/CR-6810 2003]. Separate tests of typical containment penetrations were 

conducted including tests of electrical penetration assemblies (EPAs), a personnel airlock, 

bellows, a pressure-unseating style equipment hatch, and the seals and gaskets used in 

penetrations [NUREG/CR-3855 1985, NUREG/CR-4944 1987, Clauss 1987, NUREG/CR-5096 

1988, NUREG/CR-5334 1989, NUREG/CR-5118 1989, Parks 1991, Parks, Walther et al. 1991, 

Parks and Clauss 1992, NUREG/CR-6154 1994a, NUREG/CR-6154 1995b].

In 1991, a cooperative program on containment integrity under severe accident conditions 

between the NRC and the Nuclear Power Engineering Corporation (NUPEC) of Japan was 

begun. NUPEC's funding was provided by the Agency of Natural Resources and Energy 

(ANRE) of the Ministry of Economy, Trade and Industry (METI) of Japan through a Funds-in 

Agreement between DOE and NUPEC. Testing and analyses of a steel containment vessel 

(SCV) model representative of a BWR, Mk-II containment [SAND--98-1044C 1998, SAND99-

0492C 1999] and a prestressed concrete containment vessel (PCCV) model [NUREG/CR-6810 

2003], as used in some large, dry PWR containments, were conducted. 

Efforts were also made to assess the seismic capacity of containment structures. SNL 

performed pre- and post-test analyses of shaking table tests of a 1:10-scale prestressed 

concrete containment model [NUREG/CR-6639 1999] and a 1:8-scale reinforced concrete 

containment model [NUREG CR-6707 2001]. These models were constructed and the tests 

were conducted by NUPEC at their Tadotsu Engineering Laboratory. The insights gained from 

analyzing the response of these test models were used to estimate the seismic capacities of 

typical US containments. The effects of aging-related degradation on containment capacity to 

resist severe accident pressures were investigated also [SAND98-2595C 1998, SAND2001-

1762 2001, Petti, Spencer et al. 2008]. 

1.2 Objective

All of the aforementioned research efforts are being used to set the foundation for the current 

U.S. NRC sponsored “Study on Post-Tensioning Methods.” The current study is investigating 

the structural behavior of a concrete containment vessel with grouted and ungrouted tendon 

systems, documenting and comparing in service inspection requirements provided in peer 

reviewed standards and codes for grouted and ungrouted tendons, and examining corrosion 

protection used for grouted and ungrouted tendons to assess the adequacy of the protection for 

the expected life of the structure. 

With the potential increase in use of grouted tendon systems in new plants, there is a need to 

assess the current methodology for implementing grouted tendon systems and for monitoring 

their long term structural behavior and health. This document addresses the investigation into 

the three key areas mentioned above by summarizing existing practices for inservice inspection 

and examination for PCCVs in the United States, practices in other building types, and 

international practices as well as new techniques that could potentially be used for monitoring.

1.3 The Approach

In the comparison of the structural response of grouted and ungrouted tendon systems as used 

in PCCVs, finite element method (FEM) models were created for grouted and ungrouted 

PCCVs. The 1:4 scale NUPEC/NRC PCCV tested at SNL was used as the prototype for the 

FEM model, and the results from the experimental testing were used to validate analytical

methods [NUREG/CR-6810 2003]. 

In the comparison of post-tensioning and inservice inspection methods, a comparison was 

made between different industry approaches to monitoring prestressed concrete containment 
1-3

structures. In addition, comparisons were made between different countries approaches to 

monitoring PCCVs with grouted and ungrouted tendons. A discussion is provided on the 

differences in the application of post-tensioning, and the subsequent differences in response to 

highlight the different monitoring needs of the two types of tendon systems.

For the assessment of durability and long-term corrosion protection, a combination of FEM 

analyses, probabilistic risk assessment, and review of existing literature and regulation was 

completed. 

1.4 Overview of Industry Practice for Design and Severe Accident 

Analysis

For most containment design conditions, simplified linear elastic model analyses and design 

calculations are used. This allows application of superposition principles to the treatment of load 

combinations. Further, most containment designs utilize “limit state design” as the underlying 

design philosophy by comparing calculated stress or strain values to specified limit values. This 

methodology is followed by most organizations including Électricité de France (EDF) for 

European Pressurized Reactor (EPR) containments [ETC-C 2006]. From a geometric 

standpoint, detailed analytical models are typically developed and used, but the analyses are 

primarily linear elastic. Some accounting for cracked conditions are implemented by way of 

adjustment to Young’s Moduli in various zones of the containment, and some elasto-plastic 

analyses of the liner plate (independent of the concrete) are conducted. In the U.S. the situation 

is similar for the design of containment buildings. Most designs utilize a combination of linear 

elastic finite element analysis, and simplified closed-form analytical solutions to demonstrate 

compliance with the ASME Code [ASME 2010a]. Simplified (primarily elastic) analysis 

approaches are used by some engineers while others have chosen more sophisticated 

techniques. With respect to prestressing systems, the prestressing loads are generally applied 

as fixed loads (not compliances), derived from design engineer’s in-house software. This means 

that the tendons are not modeled explicitly as part of global or local analysis models – only the 

force effects of the tendons are applied, and these sets of forces are applied once at the 

beginning of the analysis and do not change during pressurization.

In U.S. practice, there is no particular standard for analysis of containments for response to 

beyond design basis loads, although in the most recent issue of the Standard Review Plan

(NUREG-0800), there is an advisory reference to NUREG/CR-6906 for guidance on severe 

accident analysis [NUREG/CR-6906 2006] and NRC RG 1.216 and NRC RG 1.217 provide 

guidance on methods that the NRC deems acceptable for satisfying the beyond design basis 

requirements of 10 CFR Part 50. Severe accidents are typically understood as those accidents 

not included in the design or licensing basis due to their low probability of occurrence. For this 

reason, they are also referred to as “beyond design-basis” loads. As the understanding of 

severe accidents has matured, along with probabilistic safety or risk assessment methods, 

interest in the response and capacity of containment structures to resist severe accident 

conditions has also increased. Depending on national regulatory requirements and practice, it 

may be inappropriate to continue to speak of ‘design-basis’ and ‘beyond design-basis’ loads, but 

rather recognize that performance criteria (e.g. elastic stress limits versus plastic strain limits) 

will vary depending on the severity and likelihood of the loading conditions.

The goals of the analysis (‘design loads’ versus ‘beyond design basis loads’), to some degree, 

influence the requirements for, and choice of tendon modeling methodology. For severe 

accident loading conditions where the structure becomes very non-linear, modeling of the 

tendons discretely is required to achieve accurate predictions. On the other hand, for design 

level loading where the structure remains largely elastic, representing the tendons with applied 

forces can produce accurate predictions.

2-1

2 COMPARISON OF STRUCTURAL RESPONSE OF GROUTED AND 

UNGROUTED TENDON SYSTEMS

This section documents the investigation comparing and contrasting the structural behavior of a 

concrete containment vessel with grouted and ungrouted tendon systems.

As part of the investigative process, an overview of the current industry practice for PCCV 

design and severe accident analyses was completed. In addition, a comparison of existing 

methods for modeling tendons, both grouted and ungrouted, was made. Using the previously 

mentioned studies (e.g. [NUREG/CR-6809 2003]) as a starting point, finite element models 

were created of grouted and ungrouted systems. Results were presented at the OECD/Nuclear 

Energy Agency (NEA) WGIAGE Expert Meeting on Post-Tensioning Methodologies for 

Containment Building: Greased or Grouted – Consequences on Monitoring, Testing, and 

Modeling, April 20-21, 2011 in Villeurbanne, France. 

As mentioned previously, one of the first steps in this work is to review the existing analytical 

techniques for modeling tendons. New designs that incorporate grouted post tensioning 

systems in the PCCVs, require an appraisal of the current state-of-the-art practice for predicting 

PCCV response to pressure and other loads. This report will focus on finite element analysis 

(FEA) methods for modeling PCCVs and their corresponding tendon systems. To this end, the 

U.S. NRC (with the assistance of SNL and M&N) has conducted a review of existing FEA 

methods for modeling prestressing and post-tensioning, and has investigated the structural 

behavior of a concrete containment vessel with grouted and ungrouted tendons during 

prestressing and during pressure loading.

The work described herein is conducted in parallel with on-going analysis for another program –

“Standard Problem Exercise for PCCV Structural Analysis” or SPE. The structure geometry for 

both programs is the 1:4 Scale PCCV built and pressure tested to failure by Nuclear Power 

Engineering Corporation (NUPEC), U.S. NRC, and SNL in 1998-2000. A good portion of the 

experience base for FEA simulation of PCCV tendons has come from this SNL project and both 

domestic and international round robin analyses of full scale and scale model PCCVs. 

Specifically as a part of SPE 3, three FEA models have been created and Model 3 from the SPE 

3 research serves as the basis for the present investigations. Model 3, which will be discussed 

in greater detail, captures the entire 1:4 scale PCCV from the NUPEC/NRC testing. Since 

grouted and ungrouted tendon systems are practically identical up to the point of grout injection, 

the test data from the NUPEC/NRC 1:4 Scale Test, up to and including tendon anchoring, is 

relevant for both systems. Accurately modeling the tendon stress distribution resulting from the 

interplay of anchoring forces and the resisting frictional forces is important to provide the proper 

initial conditions for pressurization. Though the NUPEC/NRC test did not include a grouted 

tendon system, having test data to validate all parts of the model, particularly the tensioning and 

anchoring, is of significant value since it is test data that allows the benchmarking of 

computational models. The NUPEC/NRC test data is useful for both grouted and ungrouted 

models, because the two systems are identical during the jacking and anchoring of the tendons 

and because capturing the load distributions along the tendons that result from the tensioning 

process is critical for obtaining accurate structural response. Furthermore, the pressurization 

test data is directly applicable to the ungrouted system model which allows for additional model

refinement. 

This section summarizes methods used in the past, comments on the advantages and 

disadvantages of various methods, and also summarizes more recent work on simulating 

tendon behavior in PCCVs, as well as other civil engineering structures.
2-2

It is important to distinguish the differences between service-condition and beyond-designbasis-accident goals. From a containment pressure-response point of view, the main distinction 

is that in the service-condition loading regime, the differences in behavior between PCCV with 

grouted versus ungrouted tendons, are very small (perhaps negligible), but in the high-pressure 

regime, the differences are more significant. This is because structural response differences do 

not really manifest until containment wall deformations are large enough to overcome prestress 

and initiate hoop cracking in the concrete.

ORNL/TR 6478 and ORNL/TR 6479 document experimental work comparing the response of 

grouted versus ungrouted tendons demonstrated that grouted tendon structures develop more 

cracks with smaller widths, which results in smaller localized strains and less chance of 

penetration of corrosive elements. The testing also demonstrated that ultimate loads for grouted 

tendon beams were larger than for nongrouted beams at the same level of prestressing. Other 

conclusions from this research were that a tendon failure occurring in a non-grouted tendon is 

more critical than a corresponding failure in a grouted tendon. In addition, after first cracking, 

deflections of nongrouted beams were greater than those of companion grouted beams at the 

same load level. Grouted tendons also provide improved ultimate load capacities in flexure and 

conservatism in seating and overall anchorage efficiency. Anchorage failure is more critical for a 

nongrouted tendon than for a grouted tendon.

2.1 Comparison of Existing Methods of Modeling Grouted vs. 

Ungrouted Tendons

All further discussion in this section refers to FEM modeling that, as part of the overall suite of 

analysis activities, includes ‘beyond design basis load’ accident condition considerations. 

Stating this as an analytical goal at the outset, it becomes absolutely necessary to represent 

tendons directly in the FEA model as some kind of stiffness element, not simply as a set of 

forces. Literature review has demonstrated this to be the case.

2.1.1 Literature Review and Methods used by International Participants in 

Round-Robin Test Prediction Analysis Exercises

As previously mentioned, a good portion of the experience base for FEA simulation of PCCV 

tendons has come from Sandia participation in domestic and international Round Robin 

analyses of full and scale model PCCVs. NUREG/CR-6906 [2006] summarizes the major 

results of this and other research efforts over the last 30 years. Individual submissions on the 

part of the pre-test Round Robin analysis of the 1:4 Scale PCCV have been reviewed as part of 

the current work, published in [NUREG/CR-6685 2000]. More recently, work performed by six 

groups participating in the SPE #3 exercise has also been reviewed. In many cases, to 

complete a thorough set of predictions of behavior to overpressure, several different model 

types were developed and analyzed by each group. The modeling decisions made for the 1:4 

scale model tests are summarized in Table 2-1, and they are offered here simply as context for 

evaluating modeling choices. Also shown (for historical purposes) are modeling choices for fullscale containment analysis work performed in the late 1980s and early 1990s, sponsored by 

EPRI [EPRI NP6263-SD 1989, Dameron, Dunham et al. 1991] and in the mid-1990s by Sandia

[NUREG/CR-6433 1996]. In all known work prior to the 1:4 Scale PCCV analyses, post -

tensioning tendons were modeled as reinforcing elements or subelements.
2-3

Table 2-1 Summary of analytical models used for prior PCCV analysis studies [EPRI 

NP6263-SD 1989, NUREG/CR-6433 1996, NUREG/CR-5671 1998, NUREG/CR6685 2000, NUREG/CR-6809 2003].

Structure Scale Shape Radius/

thickness P/ Pd

Global 

models Local models Remarks

Prototypical

PWR by 

EPRI

Full

cylindrical 

concrete 

shell w/ 

steel liner & 

oblate 

spheroid 

dome

15 3.5 Axisymmetric

3D wallsegment 

model; wallbase juncture 

model

Evaluating 

prototypical 

behavior

Prototypical 

PWR by 

Sandia

Full

cylindrical 

concrete 

shell w/ 

steel liner & 

hemispheric

al dome

16.5 3.6 Axisymmetric

3D of small 

penetration. 

group

Evaluating 

prototypical 

behavior

Sizewell-B 

(PCCV) 1:10 Sizewell-B 8.6 2.4

Axisymmetric; 

3D coarse 

global (half 

symmetry)

None PRE- and 

POST-TEST

NUPEC 

(PCCV) 1:4

Large, dry 

PWR: 2-

buttress 

cylinder w/ 

hemispheric

al dome

16.5 3.2

Axisymmetric; 

3DCM

3D E/H; 

3D P/A; 

3D M/S Pen.

PRETEST

NUPEC 

(PCCV) 1:4

Large, dry 

PWR: 2-

buttress 

cylinder w/ 

hemispheric

al dome

16.5 3.6

Axisymmetric; 

3DCM; 

3D-Shell

3D E/H; 

3D M/S; 

3D tendon

ring slice; 

liner rat-hole

POSTTEST/SFMT

Examples of the FEA meshes used for the PCCV in the late 1990s are shown in Figure 2-1

through Figure 2-3. Such models (and submodels) are reasonably representative of the level of 

detail and modeling approach undertaken by many of the round-robin participants from the late 

1990s. The models were axisymmetric, and the submodels consisted of a detailed 3D model of 

the equipment hatch, and a 3D slice model of the cylinder mid-height region (called the 3DCM 

model). The model types used include: axisymmetric, 3D global, 3D sector, 2D local, and 3D 

local models of various details. Very few participants developed a full 3D model of the PCCV 

due to limitations (at that time) on computing power and on the power of model generation tools. 

As evidenced from this work and the round robin participant submittals, the minimum analytical 

assessment of containments is achieved with an axisymmetric model. (Some analysts chose to 

use a 3D sector model if the program used did not have a full suite of axisymmetric element 

types, but the analyses, and the results obtained are essentially equivalent.) Some argue that a 

3D model is also needed to capture the non-axisymmetric features of containment response. 

Setting this minimum standard depends on how the results are to be used. As discussed in the 
2-4

following sections, fully 3-D models are far more manageable to build and analyze now than 

was the case 12-15 years ago when the PCCV pre-test analysis work was performed.

The models in Figure 2-1 through Figure 2-3 included discrete tendons, which worked 

reasonably well for predicting limit states and “global” behaviors, such as radial displacements 

away from stiffness discontinuities. As described later, important limitations of these older

models were the approximations associated with simulating tendon-concrete interaction - friction 

losses and tendon-concrete relative slip.

None of the previous analysis efforts included grouted tendons. Besides needing to improve 

methods of modeling tendon-concrete interaction, friction losses, and tendon-concrete relative 

slip, there is a need for investigating methods that include grouted tendons in PCCV analyses. 

These methods are discussed in the following sections.

Figure 2-1 Axisymmetric model of 1:4-scale PCCV

Dome Tendons with 

Shell Elements

Vertical Tendons 

with Truss 

Elements

“Jacking” Element 

with Applied 

Strain

Nonlinear Springs 

at Base

Basemat
2-5

Figure 2-2 Local model of equipment hatch, 1:4-scale PCCV

Figure 2-3 Three-dimensional cylinder mid-height model (3DCM), 1:4-scale PCCV
2-6

2.1.2 Finite Element Program Modeling Methods

2.1.2.1 Basic Requirements

Containment capacity analyses aimed at predicting nonlinear response to accident loadings 

require the use of a mathematically robust, general-purpose FEA program, which has 

verification and validation documented for its intended use. Many programs meet this basic 

requirement. The selection of a particular one often depends on features that make the 

development of the model, the definition of the material stress-strain models, the execution of 

the analysis, and the post-processing of the results more efficient and convenient for the 

analyst. 

Tools for modeling concrete containments have a more recent history due to the lack in the 

early 1980s of a generally accepted concrete finite element (FE) program. The Three Mile 

Island accident highlighted the need for capability to analyze concrete containments beyond 

their design basis. An analysis tools evolving from this work are found in ABAQUS [2011], a 

general purpose FE program with special-purpose concrete and rebar subelements. The ability 

of this and other software to predict global containment response was tested and verified with 

pre-test analysis exercises performed on the Sandia 1:6 scale reinforced concrete containment 

model pressure tested to failure in 1987 [NUREG/CR-4913 1987], the Central Electricity 

Generating Board (CEGB) 1:10 scale prestressed concrete containment model pressure tested 

to failure in 1991 in the United Kingdom [SAND-90-2237C 1991], and on the Sandia 1:4 Scale 

PCCV analyses [NUREG/CR-6810 2003]. Three out of the six SPE#3 [Akin, Sircar et al. 2013]

analysis participants are currently using ABAQUS.

Similar constitutive model development efforts have occurred in the industry and have 

successfully predicted global containment model test response for the following software tools: 

DIANA, ADINA, CASTEM, NEPTUNE, NFAP, PAFEC, BOSOR5 and ANSYS. Prediction of 

global containment response has, therefore, been reasonably well established as long as 

analyses are performed with validated tools and by sufficiently experienced analysts

2.1.2.2 Specializations for Modeling PCCVs

As mentioned, the authors have performed PCCV analyses using ABAQUS. Specific 

information about the program is presented here not as an endorsement or recommendation, 

but merely to summarize important aspects of prestressed/reinforced concrete behavior 

simulation, using FEA.

ABAQUS can either be run with constitutive (stress-strain modeling) formulations found resident 

in the program (such as ABAQUS’s “damaged plasticity / smeared cracking” model) or with 

user-supplied subroutines to model complex material behaviors (such as ANACAP-U, 

developed by ANATECH). In compression, “crushing” is typically simulated as a nonlinear 

plastic behavior influenced by the presence of confining stresses. Tension and shear are 

typically modeled with concrete cracking as a “smeared-crack” formulation at the finite element 

integration points. It should be noted that “smeared-crack” is not a discrete crack formulation, 

i.e., it is not one in which nodal points are “unzipped” along cracks. Smeared crack formulations 

are the most widely used in detailed concrete FEA. The method has been compared and 

validated against many standard problems with known solutions and against laboratory 

experiments [Rashid 1968].

The steel material model used to represent rebar, prestressing, and liner should be an 

incremental plasticity model, i.e., an "incremental" theory where the mechanical strain rate is 
2-7

decomposed into an elastic and a plastic (inelastic) part. The Bauschinger Effect(1) [Woillez, 

Gingembre et al. 1977] should be simulated if cyclic response analysis is performed [Yoshida 

and Uemori 2002]. For most containment analyses reviewed by, or performed by the authors, 

the cyclic aspects of response prediction were not critical to the overall behavior, so the choice 

of hardening rules was not as important as it is for many-cycle, hysteresis problems.

2.1.2.3 Element Libraries

Based on the literature reviewed, the following recommendations apply for the choices of 

elements for modeling PCCVs:

• Concrete is typically modeled with continuum elements. In 2D planar or axisymmetry, 

8-node (quadratic edge) continuum elements are recommended, but reasonable 

accuracy can also be obtained with a finer mesh using 4-node (linear edge) continuum 

elements. In 3D, 20-node (quadratic edge) or 8-node ‘brick’ elements are recommended.

For large 3D models of the entire containment, a significant computational savings can 

be realized by using 4 node shell elements with minimal loss of computational accuracy.

• In concrete elements that crack at the finite element integration points, it is best to select 

elements with reduced integration, as long as appropriate care is taken as to mesh-size 

and to control of numerical issues such as ‘hour-glass.’ (These issues are beyond the 

scope of this report, but are usually well covered by the theory manuals of finite element 

programs.) The reason for this is improved convergence once cracking begins to 

develop.

• An axisymmetric liner should be modeled with two-node or three-node shell elements 

(whatever is compatible with the adjoining concrete elements), and beam elements used 

for representing liner angle anchors or stiffeners. In most models, except for those 

studying localized effects of liner/concrete interaction, the liner is considered to be fully 

bonded to the concrete. In 3D models, the liner should be modeled with 3D shell 

elements. 

• Smeared rebar subelements should be used to model reinforcement. Such elements 

assume strain compatibility between rebar and concrete, and have been shown to 

represent the behavior of the reinforced concrete well [Barzegar, Isenberg et al. 1993]

Smeared rebar subelements can be used to model tendons for global axisymmetric analysis of 

a PCCV. In an axisymmetric analysis, strain compatibility of hoop tendons is automatic, but with 

meridional tendons (or for all tendons in a 3D analysis), the analyst has a choice of modeling 

tendons as ordinary rebar, or as separate truss or beam elements external to the grid and 

attached to the grid with linkage or contact elements. For grouted tendons, the topic of 

simulating linkage and contact to allow sliding relative to concrete is only relevant during the 

simulation of the prestressing operation. After the structure is put into service, it is generally 

considered acceptable for the tendon elements (or subelements) to be perfectly bonded to the 

concrete. As a practical matter, however, it may be difficult to assign the correct, varying initial 

stress distribution into the tendons (accounting for losses and interaction with the concrete 

structure) without assigning discrete tendon elements which slip, relative to the concrete.

 (1) With cyclic loading, a metal’s stress/strain characteristics will change as a result of the stress distribution 

of the material on a microscopic level. This Bauschinger’s effect needs to be taken into account under such loading 

conditions. It is characterized by a reduced yield stress upon load reversal after plastic deformation has occurred 

during the initial. 
2-8

2.2 Recent (2010-2011) Developments in Tendon Modeling

2.2.1 Special Consideration of Anchorage Zones

Behavior at and near anchor zones is worthy of further discussion, and may even require 

consideration for further analysis modeling detail. For tendons, the anchorage zones are an 

important design detailing problem; structural analysts can well assume that tendon anchorage 

hardware will perform as specified, and in most cases, the diagonal splitting forces in the 

concrete which surround the anchorages are addressed well by reinforcing design codes

[Collins and Mitchell 1997, ASME 2010a]. In a complete analytical review of the structure, 

however, these zones should not be ignored completely. Sometimes anchor zone stresses in 

the concrete can produce unwanted cracking when the stresses combine with structural loading 

stresses (shear and flexure), so these zones should be included on structural evaluation 

checklists. If such stress combinations become questionable, a local FE model can be used to 

evaluate them, and in such cases, the tendons should be modeled as independent straining 

elements first, then have strain compatibility “locked in” as a secondary stage of analysis as has 

been demonstrated in this report. This is the only way to accurately represent localized stresses 

that occur in the concrete in the immediate vicinity of anchorages. 

In terms of comparison of ungrouted to grouted behavior, it should be noted that the stressed 

condition in the anchor zones is of long-term performance concern for ungrouted tendon 

systems, but is less so for grouted tendons, assuming high quality grout placement. Highly 

stressed concrete can continue to creep, and micro-cracks which can turn into full-fledged 

cracks over decades of structure service. In the anchor zones, this phenomenon is of somewhat 

less concern for grouted tendons than ungrouted tendons. 

For historical reference, there was a PCCV scale model test in which anchorage zone stresses 

proved to be very significant to the prediction of the response and failure mode of the structure. 

This case is described by the following excerpts from [NUREG/CR-5671 1998].

“In July, 1989 a test was conducted for a 1:10-scale model of a prestressed concrete 

containment vessel by the Central Electricity Generating Board (CEGB) in the United 

Kingdom. This model of the Sizewell-B NPP containment structure utilized seven-wire 

strands in plastic sheaths to represent the vertical hairpin and hoop tendons. The vertical 

tendons were anchored below the basemat, and the 240 ̊ hoop tendons were anchored 

in two of the three vertical buttresses. The model was unlined (although a bladder was 

inserted in the model to prevent leakage) and included an equipment hatch penetration. 

The model was tested hydrostatically. The NRC, through an agreement with the United 

Kingdom Atomic Energy Authority (UKAEA), participated in the test program with SNL 

and ANATECH Corporation providing technical support to the NRC.

The geometry is shown in Figure 2-4. The model was tested with four loading cycles, 

each from zero pressure to 1.15 x design pressure, followed by a single ultimate 

pressure test. Design pressure of the model was 0.345 MPa (50 psig). During the high 

pressure test, a maximum pressure of 0.834 MPa (121 psig) was reached at the model 

base (2.4 x design pressure). Failure occurred when bending of the basemat slab led to 

rounding of the underside of the basemat, model tilting and potential instability, spalling 

of basemat under-surface concrete, and termination of the test. Upon post-test 

inspection, model failure was found to have been associated with the basemat spallation 

and the resulting loss of bond in basemat bottom reinforcement as illustrated in the 

sketch of Figure 2-5. The cracking and spalling which led to the local failure was 

adjacent to the vertical tendon anchorage zone, and the anchorage zone stresses 

exacerbated the cracking. Another interesting result was the observation was made 
2-9

regarding the post tensioning system for the scaled model. Due to scaling difficulties, it 

was decided to use a single greased strand within a narrow plastic sheath as opposed to 

a traditional metal duct. Despite the presence of grease, this configuration showed 

evidence of significant friction, and later, by parametric analysis, the level of friction was 

shown to influence model behavior. Evidence of this is shown in Figure 2-6.

Figure 2-4 Schematic of 1:10 scale Sizewell model [SAND-90-2237C 1991]
2-10

Figure 2-5 Section view of posttest condition of the 1:10-scale model 

[NUREG/CR-5671 1998]

Several analyses versus experiment comparisons are shown in Figure 2-6 where the 

“Bonded Tendon” and “Unbonded Tendon” curves represent axisymmetric modeling 

results and the “Gauge” curves refer to instrumented results from the experiment 

described in [SAND-90-2237C 1991]. As with a number of other experiments, 

axisymmetric modeling resulted in satisfactory prediction of containment cylinder global 

response, but the modeling was unsatisfactory for predicting local and 3D effects such 

as rebar de-bonding, localized concrete spallation, and behavior near penetrations.”
2

-11

b

d

Figure 2-6 Comparison of pretest analyses vs. experiment results for 1:10 scale P/C model [SAND-90-2237C 1991]. 

a) Radial displacements at midheight, b) Basemat uplift displacement, c) Hoop tendon strain at midheight, d) Meridional rebar strains. 

ac
2-12

The parametric analysis showed that, especially for basemat uplift behavior, and 

characterization of meridional bending at the base of the wall, a grouted tendon representation 

simulated the structure behavior much better than an ungrouted tendon representation. From 

this and other observations made on strain gradients occurring in the hoop tendons, a final 

conclusion was drawn that plastic-covered tendons behave more like grouted tendons than 

ungrouted. To physically simulate ungrouted tendon behavior in scale model tests requires use 

of tendon ducts (sheaths) which allow free-sliding of tendons within the ducts.

2.2.2 Comments on State-of-the-Practice for Modeling Tendons in Other 

Structures

Because nearly all PCCV’s in service in the United States have ungrouted tendons, there is little 

design experience and long-term inservice observation experience existing in the United States 

Nuclear Industry (both regulatory and industry side). Therefore, it is beneficial to look elsewhere 

to gain this design experience and inservice observation experience. Some examples of 

external industry experience include: (a) post-tensioned box-girder bridges (these are almost 

always grouted, and California has more of them than any other state) (b) buildings (not always 

grouted, so bridge performance is probably more relevant), and (c) overseas nuclear experience 

(still developing). The following discussion of a box-girder study is a valuable starting point for 

ensuring the capability of accurately modeling grouted tendons.

For building and bridge structures, it is common to consider the effects of prestressing by only 

applying forces to the structures, and not including tendons as stiffness elements. A variety of 

commercial software exists which calculates tendon losses of all types described herein, and 

applies the distributed effects of the tendons directly to the structure. Unlike for containments, in 

post-tensioned bridges, the tendon forces can be the result of tendons along curved paths 

(parabolic or other “draped” tendon designs), which by design, transfer vertical load countering 

forces and moments to the bridge. 

In recent years, especially in cases where prestressing systems have exhibited service 

problems (such as unexpected concrete cracking associated with the prestressing), more 

sophisticated tendon modeling techniques are employed to study the local concrete stresses in 

the vicinity of the tendons [NCHRP 2009]. Examples of such modeling are shown in Figure 2-7

through Figure 2-9. In the first example, the physical problem studied is the tendency for lateral 

pull-out from curved box-girder webs; in this case, the tendon ducts and lateral forces caused by 

the tendons are modeled in detail (though the tendons themselves are not modeled). In the 

second example, the physical problem studied is cracking near the jacking/anchorage zone of 

tendons in post-tensioned box girders. In these studies, the tendon wires are modeled as 

discrete elements (as in PCCV models), and various methods for allowing tendons to slide in a 

tensioning step, then become bonded in a “grouting” step were studied. These studies [NCHRP 

2009] resulted in the development and use of the “slot” methodology that was used recently in 

the PCCV studies (subsection 2.3.3). These FEM models provided validation of methodology 

used for the PCCV grouted modeling described in this report.
2-13

Figure 2-7 Study of tendon pull-out problem in curved concrete post-tensioned box 

girders [NCHRP 2009]
2-14

Figure 2-8 Study of tendon anchorage zone cracking, view of typical FE mesh [NCHRP 

2009]
2-15

Figure 2-9 Study of tendon anchorage zone cracking, view of tendon mesh [NCHRP 

2009]

2.2.3 Other Issues for Consideration – Reduced Bond of Tendons Compared to 

Reinforcement

European Codes require the rebar in containments to be anchored/bonded in order to develop 

the full capacity of the rebar, but for the tendons the rules are different. That is, the initial anchor 

forces, prior to grouting are limited to less than full tensile capacity of the tendon, and as a 

result, after grouting, the combination of the bond stress and the anchor force are not 

necessarily adequate to develop the full capacity of the tendon. The experience from the tests of 

the 1:4-scale PCCV model suggests, however, that the anchorage was adequate to develop the
2-16

full tensile capacity of the tendons, even though the tendons in this model were ungrouted. This 

was most likely due to the margins in the anchorage design rather than any conscious effort to 

design the anchors for the full tendon capacity.

There are also issues related to consideration of flexural cross-sections. In design, some codes 

allow counting the tendon as a steel element in the flexural resistance of the cross-section, and 

some codes do not. ASME BPVC Section III, Division 2 (CC-3541(b)(3)) [ASME 2010a] does 

allow grouted tendons to be included in the cross section for flexural design, while the 

Eurocodes do not. At sections with large flexure demand (large curvature) the tendon cannot be 

counted on to strain perfectly compatible with the concrete, like a rebar does. Since the tendon 

typically is located near the neutral axis, its contribution to flexural strength may be limited; 

however, for design in high flexural regions, it might be appropriate to consider the section both 

ways (with or without the tendon contribution) although the codes do not require this.

From an analysis perspective, the question regarding tendon bond is relevant especially for 

severe accident analysis. Laboratory tests referenced by Collins and Mitchell [Collins and 

Mitchell 1997] have shown that tendons can develop full strength within a remarkably short 

development length, on the order of 1 meter (3.3 ft.) or less. Thus it is generally recommended 

that for grouted tendon systems, FE simulation of the tendons use a fully bonded assumption as 

would normally be done for reinforcing bars.

2.3 Finite Element Models of PCCV with Ungrouted and Grouted

Tendons

2.3.1 FE Representation of Rebar

The modeling of rebar varies among FE programs. Generally for large, complex structures such 

as PCCVs, displacement compatibilities between the rebar and the concrete must be assumed. 

Numerically, this means that the rebar is represented by the same displacement shape 

functions as the continuum element in which they reside. This is not universally observed in all 

programs. Another challenge is the translation of rebar drawings to computer input. For 2D 

axisymmetric grids, rebar modeling is relatively straightforward. However, for 3D continuum 

analysis, the rebar arrangements are very complex, and it is virtually impossible to develop input 

without preprocessing software. ABAQUS has a built in function which allows the user to define 

rebar layers, and embed them in a concrete section. The user must enter the cross-sectional 

area of the rebar, the spacing, and the material for the rebar. The rebar layer is then defined as 

embedded in the concrete to ensure displacement compatibilities between the rebar and the 

concrete. This method was used to define the rebar in the grouted and ungrouted models.

2.3.2 FE Representation of Tendons

In PCCVs, prestressing is provided by an arrangement of steel tendons. Typically, each 

horizontal tendon makes a complete loop of the containment and is anchored within a buttress. 

Each horizontal tendon is tensioned on both ends. The vertical tendons can be of two different 

types: curved (sometimes called ‘hairpin’ tendons), and ‘pure’ vertical tendons. The curved 

tendons are vertical tendons which are returned through the dome and tensioned at both ends. 

For straight vertical tendons, the upper end is anchored at a dome ring and the lower end is 

anchored in the vertical tendons pre-stressing gallery, located underneath the support slab. 

Tendons are located in steel ducts, or sheaths (see Figure 2-10 and Figure 2-11). The 

connections between sheaths consist of sleeves with a length of 4 times the sheath diameter; 

both ends being sealed by a thermo-retractable sleeve. As stated previously, for severe 
2-17

accident analyses, the tendons should be modeled as separate stiffness elements (as opposed 

to an applied force). This is typically accomplished with beam or truss elements and the use of 

both element types has precedence (see NUREG/CR 6809 and NEA/CSNI/R(2005)5

respectively). Since the cross-sectional properties of the tendons are small in comparison to the 

containment wall, the contribution to section moment capacity is primarily derived from the axial 

behavior of the tendons. For modeling the interaction between the tendon and the ducts in 

commercial FEA codes, the use of beam elements may allow more options than truss elements.

The tendon connection to concrete can be accomplished by modeling the tendon ducts 

separately or by using connector elements to transfer load from the tendons to the concrete. For 

either case, capturing the friction between the tendon and the tendon duct is important for 

achieving accurate results. 

Figure 2-10 Tendon ducts in 1:4-scale PCCV (silver colored tubes) along with 

traditional rebar 
2-18

Figure 2-11 Tendon ducts (silver colored tubes) in 1:4-scale PCCV with and without 

connections as well as traditional rebar

The tensioning of horizontal and vertical tendons has a specific sequence which must be 

followed in order to avoid excessive flexural effects. The tensioning of the cables is typically 

carried out when concreting of the dome has been completed and when the concrete of the last 

layer has aged more than 28 days such that sufficient compressive strength has been attained 

to ensure that the concrete has sufficient mechanical resistance to support the tensioning of the 

cables and thus limit the concrete creep deformations.

For grouted tendons, following tensioning, the ducts are injected with a cementitious grout, the 

intention of which is to completely fill the voids between the tendons and the duct walls. There 

are time limits placed on the tensioning operation and the grouting operation to ensure a limited 

exposure of unprotected tendons to potentially deleterious environmental conditions. 

2.3.3 Simulating Initial Conditions and Losses

Prestressing losses should be estimated to accurately represent the actual stresses that will 

exist in the containment. In general, the philosophy used should be:

1) calculate best estimate “in service” values based on the nominal design values, modified 

for creep or any other in situ conditions

2) apply tendon stresses according to best estimate values, and allow the model to 

equilibrate to final tendon stresses which are reasonably close to best estimate values, 

including anchor slip

In axisymmetric analysis, there is no opportunity to simulate progression of friction along the 

tendon path in the hoop tendons, but this phenomenon can be included in the meridional 
2-19

tendons. Standard prestressing losses, along with a brief explanation of the basis for their 

consideration are listed below [Regulatory Guide 1.35.1 1990]:

Instantaneous or Initial Losses

1) Elastic Shortening – occurs simply due to equilibrium-seeking displacement within the 

finite element analysis

2) Anchorage Slip – can be considered explicitly in local model analysis; but not directly 

relevant for axisymmetric analysis. For conventionally anchored tendons, anchor slip in 

the range of 4-6mm (0.16-0.24 in.) is typical. The amount of actual stress loss resulting 

from this is tied to the angular friction assumption

3) Angular Friction (and wobble friction) - Considered explicitly in local analysis of hoop 

tendons, and in global axisymmetric analysis of meridional tendons in dome. Also 

considered in calculating the average hoop tendon stress to be assigned in global 

axisymmetric analysis (Also see below)

Time Dependent Losses

4) Steel Relaxation – normally considered by the containment designers in calculating the 

"nominal inservice" values

5) Shrinkage of Concrete – normally considered in combination with creep, and also, 

normally considered by designers in calculating “inservice” values

6) Creep of Concrete – studied in detail for the PCCV, but for design, should probably be 

addressed using standard design formulae. The combination of creep and shrinkage 

comprises a significant prestress loss over time, often reaching 5%-10% of overall 

prestressing

7) Others: Temperature – generally not considered, but it should be noted that for a thermal 

analysis, if tendons become significantly heated, significant loss of prestress can occur 

simply due to thermal expansion and the associated stress relaxation

With post-tensioning, the amount and distribution of elastic shortening depends on the order of 

post-tensioning. In general, it should be assumed that the tendons are jacked in a sequence 

appropriate to reacting the total desired lock-off force. In ABAQUS, an option called 

"PRESTRESS HOLD" allows an initial post-tensioning equilibrium step that holds the tendon 

stresses at a preset value while the structure iterates to equilibrium and thus maintains a 

constant stress regardless of elastic shortening. This can generally be achieved without using 

the PRESTRESS HOLD option, within a few attempts, by applying larger than target prestress, 

reaching equilibrium in the containment, noting the final prestress, then adjusting and applying 

the prestress load again. This iterative procedure is used in the current analysis work.

Angular friction, wobble friction, and friction in straight portions of meridional tendons in the 

cylinder below the springline can generally be neglected. From standard prestressed concrete 

texts [Collins and Mitchell 1997], the angular friction should be included in the curved tendon 

portions with the formula:

( )

2 1

k x T Te− + α μ = (2-1)
2-20

Where:

• α is the angle between T1 and T2

• μ is the coefficient of static angular friction

• k is the coefficient for wobble friction

• x is the length along the tendon

• T1 is the tendon force next to a jack before friction losses

• T2 is the tendon force at some angle α away from T1.

It should be noted that wobble friction, while often estimated in design calculations, is not 

included in modeling, and will be neglected moving forward in the report, since the tendon 

follows a smooth arc. In the 1:4 Scale PCCV, the angular friction coefficient was observed (by 

ancillary testing) to be 0.21, but for a full-scale prototype, a coefficient of 0.15 is generally more 

appropriate. This difference is primarily attributable to the smaller radius of the scaled 

containment model in comparison to a full-scale containment and a small contribution from the 

wobble component that is not separately considered. 

For example, consider a meridional tendon with α = 90° = 1.57 Radians, and μ = 0.21 (from 

specifications),

(2-2)

. (2-3)

Therefore, the percentage loss from the springline up to the dome apex is approximately 7.6 %.

One of the conclusions of the 1:4 scale PCCV project was determining the importance of tendon 

friction modeling, and should therefore be addressed. The following are various ways of 

capturing this analytically, in decreasing order of complexity: 

1) An advanced contact friction surface between the tendons and the concrete, 

2) Pre-set friction ties (as shown in Figure 2-12, used in the 1:4 Scale pretest/post-test 

analyses [NUREG/CR-6809 2003]) 

3) If neither of these methods are practical within the scope of the calculation, it is best to 

start with an “average” stress level (using a friction loss design formula), but assume 

uniform stress distribution in the tendons throughout pressurization, i.e., an ungrouted

tendon assumption

4) Same as 3, but using a grouted tendon assumption 

The drawbacks of the pre-set methods (2 through 4 above) can be noted by comparing the FEA 

predicted tendon stress profiles (black, blue, and orange curves of Figure 2-13) to the tendon 

stress measurements from the test (symbols shown in Figure 2-13). Figure 2-13 and similar 

plots of other hoop tendon behavior show that while the tendon stress distribution, after 

prestressing, follows closely to that predicted by a Method 2 FE Model (following design 

equations), once larger magnitudes of pressures are applied, the tendon stress profile becomes 

much different – by high pressure (first tendon yield and beyond), the tendon stress profiles 

become quite uniform. This observation became one of the drivers of the current development 

work of the standard problem exercise (SPE) #3.

( ) 1.57 0.21

1 2 2 1.0817 x T Te T − = =

2 1 T T = 92.4% 
2-21

Figure 2-12 Method of simulating tendon friction using friction truss-ties
2-22

Figure 2-13 H53 tendon force comparisons to pretest (From NUPEC/NRC PCCV test 

at SNL)

During the analysis work performed for SPE #3, more advanced representations for ungrouted

tendons have been developed. A simple model (Model 1, Figure 2-14 and Figure 2-15) 

representing a ring-slice through the cylinder (for the 1:4 Scale PCCV) was the test bed for this 

development.

Figure 2-14 Model-1 Abaqus Model

150000

200000

250000

300000

350000

400000

450000

500000

550000

600000

650000

Force (newtons)

Azimuth (degrees)

H35 Tendon Force Distribution, El. 4572

(Load Cells and Average of Wire Strain Gages) Design @ Max

Tension

Design After

Seating

@ Max

Tension

After Seating

0 x Pd

1 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

2.6 x Pd

2.0 x Pd

0.9 x Pd

0 x Pd

Analysis at

Anchor Set

Analysis at 3 x

270° 360, 0° 90° 180° 270° Pd

0° 180°

270°

90°

Tendon H53

Elevation: 6.579m 

Anchor: 270°

270°

F/W

M/S

A/L

E/H

90°

270°
2-23

Figure 2-15 Tendon layout

Figure 2-16 shows the tendon connectivity to the buttresses. For this model the tendons are 

placed in contact with concrete using a “contact surface,” and friction equal to 0.21 is assigned, 

based on the ancillary testing performed for the 1:4 Scale PCCV. Prestress is prescribed by 

only applying initial stress to the single tendon element outside the concrete mesh and allowing 

the FE solution to reach equilibrium. The prestress (prior to pressure load application) is 

reached in two solution steps:

1) Stress is applied to the tendon ends (call these the jacking elements); this produces the 

dark blue curve in Figure 2-17.

2) Relaxing the stress in these ends by amount equal to the “anchor slip” (3.95 mm (0.156 

in.)); this produces the red curve in Figure 2-17. Tendon strains are shown in Figure 

2-18.
2-24

Figure 2-16 Anchorage of tendon to concrete

Figure 2-17 Tendon stress
2-25

Figure 2-18 Tendon strain

Even more refined development of the ungrouted tendon simulation was developed for an 

Equipment Hatch Model (“Model 2” in the SPE 3 work). Tendon beam elements were placed in 

metal “sheaths” or “ducts” characterized by shell elements, followed by the shell elements 

embedded in the concrete. This method was demonstrated to work, but this method still needs 

to be relegated to local model analysis due to the computational expense. For global models of 

entire PCCVs, the method demonstrated in Model 1, using contact algorithms (but no explicit 

modeling of sheaths) is the most practical. 

With the methodology followed in Model 1, this contact condition requires that the nodes of the 

tendon and the nodes of the concrete be coincident. With the complexity of the tendon 

geometry, making the concrete mesh compatible with the tendon mesh is extremely difficult and 

time consuming. Therefore, an innovative strategy was developed to facilitate the modeling of 

the tendon-concrete interaction. This novel technique captures the interaction of the tendons 

with the reinforced concrete structure and includes the associated frictional effects. Every node 

of the tendons has a matching reference node that shares the same space. These reference 

nodes are tied to the surface of the concrete and transfer forces and displacements directly to 

the concrete. Connector elements are used to constrain the tendon nodes to the reference 

nodes. ABAQUS provides a selection of connector types, and the SLOT connector elements 

have been selected. SLOT connectors, as the name implies, allow the tendon nodes to move 

only in one direction relative to the reference node. This direction is assigned to be the initial 

tangential direction along the tendon. The connector elements are able to solve for the frictional 

resistance by taking the force normal to the direction of motion and determine whether sticking 

or slipping occurs. The traction and normal forces exerted by the tendons are transferred 

directly to the concrete through the reference nodes.

This novel approach is conceptually equivalent to the use of contact surfaces, as described 

previously. To compare the two approaches numerically, Model 1 was modified to use the new 

0.000

0.005

0.010

0.015

0.020

0.025

-360 -315 -270 -225 -180 -135 -90 -45 0 45 90 135 180

Strain

Azimuth (degrees)

3.66xPd

3.4xPd

3.3xPd

3.0xPd

2.5xPd

2.0xPd

1.5xPd

1.0xPd

Anchor

Tension

Tendon A -90 0 90 135 180 225 270 315 360 45 90 135 180

Tendon B -180 -135 -90 -45 0 45 90 135 180 225 270 315 360
2-26

approach and the results obtained are plotted in Figure 2-19 and Figure 2-20. These figures 

compare the two approaches on the basis of strain in the tendons and the deformed shape of 

the PCCV. Both methods compare well, with the observed difference that the distribution in 

stress/strain in the tendons is not as smooth with the slot connectors as with the contact surface 

approach.

a

b

Figure 2-19 Tendon strain with (a) contact surfaces and (b) connector slots used to 

define contact

0.003

0.004

0.005

0.006

0.007

0.008

0.009

0.010

-360 -315 -270 -225 -180 -135 -90 -45 0 45 90 135 180

Strain

Azimuth (degrees)

Tendon Strain - Contact Surface

3.3xPd

3.0xPd

2.5xPd

2.0xPd

1.5xPd

1.0xPd

Anchor

Tension

Tendon A -90 0 90 135 180 225 270 315 360 45 90 135 180

Tendon B -180 -135 -90 -45 0 45 90 135 180 225 270 315 360

0.003

0.004

0.005

0.006

0.007

0.008

0.009

0.010

-360 -315 -270 -225 -180 -135 -90 -45 0 45 90 135 180

Strain

Azimuth (degrees)

Tendon Strain - Connector Slots

3.3xPd

3.0xPd

2.5xPd

2.0xPd

1.5xPd

1.0xPd

Anchor

Tension

Tendon A -90 0 90 135 180 225 270 315 360 45 90 135 180

Tendon B -180 -135 -90 -45 0 45 90 135 180 225 270 315 360
2-27

a

b

Figure 2-20 Deformed shape with (a) contact surfaces and (b) connector slots used to 

define contact

-10,000

-8,000

-6,000

-4,000

-2,000

0

2,000

4,000

6,000

8,000

10,000

-10,000 -8,000 -6,000 -4,000 -2,000 0 2,000 4,000 6,000 8,000 10,000

mm

mm

Model 1 - Contact Surface disp amp = 100x

Tension

0.0xPd

1.0xPd

1.5xPd

2.0xPd

2.5xPd

3.0xPd

3.24xPd

-10,000

-8,000

-6,000

-4,000

-2,000

0

2,000

4,000

6,000

8,000

10,000

-10,000 -8,000 -6,000 -4,000 -2,000 0 2,000 4,000 6,000 8,000 10,000

mm

mm

Model 1 - Connector Slots disp amp = 100x

Tension

0.0xPd

1.0xPd

1.5xPd

2.0xPd

2.5xPd

3.0xPd

3.24xPd
2-28

The advantage of using slot connector elements over the contact surfaces used in Models 1 and 

2 is the simplicity of the interaction. Although each connector adds to the number of elements in 

the model, it is much less than the elements needed to make tendon ducts as in Model 2. With

simplicity come limitations, however. As stated above, the tendon nodes can only move in one 

direction that is assigned before the analysis begins. As the tendons slip and move, the line of 

motion does not. With the contact surfaces, the interaction between the tendon and ducts 

adjusts for the new position the nodes are in for each analysis increment. For detailed models 

focusing on a small, local region, using contact surfaces is the desirable method to use, and is 

feasible. The slot-connector approach is more appropriate for full 3D global models.

Another behavior which can be analyzed using Model 1 is the tendon slip relative to the 

concrete. This is shown on Figure 2-21 and Figure 2-22. This shows graphically how the tendon 

travels over 70 mm (2.76 in.) at the jacking points before it is anchored Figure 2-21, but at 

exactly halfway between the jacking points, the tendon travel is zero. Figure 2-22 shows the 

amount of slip that occurs during pressurization up to near failure pressure (notice the smaller 

range on the abscissa). Though relatively small, this nearly 3 mm (0.118 in.) of slip during 

pressurization (Figure 2-22) is important for allowing the tendon stresses to redistribute (for 

ungrouted tendons) during pressure loading. This is why the curves appear to be coincident in 

Figure 2-21. These slippage amounts (after prestressing) are all zero for the grouted tendon 

case. 

Figure 2-21 Tendon slip vs. location resulting from jacking the tendon from the unstressed state (ungrouted)

-100

-80

-60

-40

-20

0

20

40

60

80

100

-315 -270 -225 -180 -135 -90 -45 0 45 90 135

Tangential Slip (mm)

Azimuth (degrees)

Anchorage

1 Pd

2 Pd

3 Pd

3.4 Pd

Tendon A 45 90 135 180 225 270 315 360 45 90 135 

180 Tendon B 225 270 315 0 45 90 135 180 225 270 315 
2-29

Figure 2-22 Tendon slip vs. location after anchorage, during pressurization (ungrouted)

2.3.4 Material / Failure Criterion

Tensile properties of prestressing determined for quality control usually test the individual wires 

or strands rather than full tendons. Sandia’s experiments suggest that the tensile properties of 

the tendon system, including anchor hardware are weaker and less stiff than would be expected 

from the strand test data [NUREG/CR-6809 2003]. Since it is not always feasible to test the 

tendon systems to failure, a reduction factor (5 to 10%) should be applied on stiffness and 

strength of the individual wire or strand data. On stiffness, this reduction is associated with 

“strand wrap angle” effects. On strength, this reduction is associated with failures in the 

anchorage systems. It should be noted, however, that it is possible, but not 100% certain, for 

tendon systems to achieve their full strength as implied from a strand test.

For PCCV structural analysis in which liner tearing and leakage is not the focus of the analysis, 

the relevant failure criterion is tendon failure. The rebar generally has substantially higher 

ductility than tendons, so rebar are not the controlling criteria. Tendon failure criteria should 

come from ancillary tests of the full tendon system, not just from strand tests. For the PCCV, 

such tests resulted in a tested ductility limit of 3.8% strain, as shown in Figure 2-23. This is used 

directly as the failure criterion for FEA. The ductility and strength of the tendon-strand and wire 

material has not changed much in the past 20-30 years, but the jacking and anchoring hardware 

has been improved. The 3.8% ductility was for the tendon/strand acting alone, irrespective of 

the jacking and anchoring hardware. There were a few outliers in the PCCV test data showing 

lower ductility, most likely caused by failures at or near the hardware. 

If liner failure is included in the prediction, this will very likely occur prior to tendon failure. The 

steel material model used for FEA of PCCVs is an incremental plasticity model (i.e., where the 

mechanical strain rate is decomposed into elastic and plastic (inelastic) parts). A key 

characteristic in this formulation is the yield surface, which is used to determine if the material 

responds purely elastically at a particular state of stress. We are using the Mises yield surface 

to define isotropic yielding. This is appropriate for all but the most extreme scenarios, e.g., sheet 

metal forming or welding, where the Hill yield surface (anisotropic yielding) or a porous metal 

plasticity model may be more appropriate.

-4.0

-3.0

-2.0

-1.0

0.0

1.0

2.0

3.0

4.0

-315 -270 -225 -180 -135 -90 -45 0 45 90 135

Tangential Slip after Anchorage (mm)

Azimuth (degrees)

Anchorage

1 Pd

2 Pd

3 Pd

3.4 Pd

Tendon A 45 90 135 180 225 270 315 360 45 90 135 

180 Tendon B 225 270 315 0 45 90 135 180 225 270 315 
2-30

The present research uses an isotropic hardening law (i.e., the yield surface changes size 

uniformly in all directions as plastic straining occurs). A stress vs. strain curve is input to the 

program, but how it is implemented is not uniaxial stress vs. uniaxial strain, rather it is effective 

(Von Mises) stress versus effective plastic strain. The yield surface grows (or contracts) as a 

function of the effective plastic strain, and can shift the relationships of principal stresses and 

thereby influence strain distribution after yielding occurs.(2)

 

It is important to note that even with sophisticated plasticity models, failure is typically predicted 

externally by the analyst, applying a strain failure criterion which takes into account the triaxiality 

of the stress state. We do this using the Davis Triaxiality Factor defined by the formulas shown 

below, and which have been the failure criteria of choice in nuclear containment analyses for 

many years. 

Biaxial-stress based Failure Criteria

(2-4)

Where μ is the ductility (reduction) ratio and TF is the Davis Triaxiality factor

(2-5)

When the third principal stress is zero or nearly zero, as in the case of the shell theory,

(2-6)

For instance when σ1 = σ2, TF = 2 and the ductility ratio is 0.5 (i.e., failure strain reduces to half 

its uniaxial value). For the last two decades, many containment analysts have used this criterion 

for predicting onset of liner tearing, but most have concluded that there is also extensive 

judgment involved in its application. Strains predicted by FE models can be highly dependent on 

the level of detail (and mesh refinement) included in the model. And, as was seen in the 1:4 

Scale PCCV Model, the existence of flaws in the material (especially at weld seams) mean that 

tears might occur at strains significantly lower than the absolute ductility of the material. 

 (2) ABAQUS provides an isotropic hardening model, which is useful for cases involving gross plastic straining 

or in cases where the straining at each point is essentially in the same direction in strain space throughout the 

analysis. Isotropic hardening plasticity is discussed in more detail in “Isotropic elasto-plasticity,” Section 4.3.2 of the 

Abaqus Theory Manual "Abaqus 6.11" (2011a). Dassault Systèmes Simulia Corp., Providence, RI.

1 2 TF μ − =

( )

( ) ( ) ( )

123

1

2 22 2

12 23 31

2

TF

σσσ

σσ σσ σσ

+ + =

  − +− +−  

( )

1 2

1

2 2 2

1 12 2

TF σ σ

σ σσ σ

+ =

− +
2-31

Figure 2-23 Tendon stress strain curve

2.4 Three Dimensional Global Finite Element Model

For the calculations presented in this work, the geometry is as specified in the SPE3 problem 

statement, and on the 1:4 Scale PCCV drawings, shown immediately below in Figure 2-24. The 

model used for the present study is based on “Model 3” from the SPE 3 analyses [Akin, Sircar et 

al. 2013] and the term Model 3 will also be used here. 

Figure 2-24 The geometry for model 3 is based on the 1:4 scale PCCV test (Figure 1.2 

from [NUREG/CR-6810 2003])
2-32

Simulation of friction and pressure response related changes to tendon stress distribution are 

included in the analysis. The model includes concrete, tendons (hoop and vertical), rebar (hoop, 

vertical), and liner. Shear (through thickness) reinforcement is not included, since the structure 

wall is represented by shell elements, which fundamentally do not calculate any throughthickness stress. Vertical tendons and rebars are included. Concrete is modeled with 4-node 

shell elements, and rebar is modeled with embedded subelements. Tendons are modeled with 

two-node beam elements, and the liner is modeled with with 4-node shell elements that 

consider 5 integration points through the thickness, overlain onto the same nodes as the 

concrete shell nodes, but offset by the appropriate eccentric dimension. The concrete basemat 

rests on a set of non-linear springs that simulate the deformation characteristics of the soil and 

allow for basemat uplift. Additionally, the material constitutive behavior used was obtained from 

the 1:4 scale PCCV testing and represents as-tested material properties as opposed to design 

values [NUREG/CR-6810 2003]. 

Analytical representation of initial losses was addressed naturally within ABAQUS through the 

use of friction, as described in the previous section. Every tendon was modeled, and each 

tendon had a “jacking element” (similar to the methodologies developed in Model 1 and Model 

2) protruding from the tendon end zone.

The geometry for Model 3 is shown in Figure 2-25 through Figure 2-27. Figure 2-25 shows the 

general outline of the Model 3 FE Mesh. The wall-base juncture occurs at the correct location, 

geometrically, but since shell elements are aligned with mid-thicknesses of structural elements, 

wall-base juncture is separated by half the thickness of the basemat. This juncture is 

appropriately tied with translational and rotational constraints (“rigid links” as shown in Figure 

2-28). Figure 2-26 and Figure 2-27 show the actual element mesh, including color coding of 

different rebar mesh densities.

Figure 2-25 Model 3 overview
2-33

Figure 2-26 Meshed concrete vessel with various section assignments. Variations due 

to rebar layers and concrete thickness. Thickness of shell element 

rendered in ABAQUS

Figure 2-27 Meshed concrete vessel with various section assignments. View of M/S and 

F/W
2-34

Figure 2-28 Rigid links from bottom of vessel to basemat elements

Figure 2-29 and Figure 2-30 show the tendon “jacking elements” (or “nubs”) employed in the 

model. These elements have elastic properties, and all of the prestressing is applied through 

these elements by applying temperature contraction to them. Then the stresses distribute all the 

way around the vessel in the tendons (during solution equilibrium iterations), similar to how ‘real 

world’ tendons are stressed. During the prestressing step, the ends of the jacking elements are 

rigid-linked to the nearest concrete node. The analysis sequence simulation for tendon jacking 

is illustrated in Figure 2-31 through Figure 2-33, and described below.

Figure 2-29 Vertical tendon jacking element ends rigid linked to closest basemat node
2-35

Figure 2-30 Hoop tendon jacking element ends tied to closest buttress center node
2-36

Figure 2-31 Hoop tendon jacking elements (‘nubs’) and tendon nodes shown relative to 

concrete nodes in the buttress region

As shown in Figure 2-31 and Figure 2-32, the outer ends of the jacking elements “nubs” are 

mathematically “tied” to a concrete node in the center of the buttress; the limitations of shell 

element representation of the buttress zone caused problems and unrealistic behavior when 

these nodes were tied to the exterior buttress nodes (because in fact, no node exists at the 

exact, exterior jack location). A more accurate model of the anchoring behavior could be 

obtained with hex elements, however this increase in accuracy would come with a great 

increase in computational expense owing to the increased element count. Anchoring the 

tendons in the centroid is sufficiently accurate to predict global behavior with a high degree of 

confidence. The next tendon nodes moving inward from the ends are (1) located at the center of 

the buttress (when the tendon is in the undeformed position), and (2) at the tendon “tangent” 

point, i.e., when the tendon (as-built in the structure) begins its curvature around the concrete 

cylinder. This node is the first point where the SLOT constraints with friction begin;
2-37

Figure 2-32 Deformed shape of hoop tendon anchor system after “jacking loading 

step”

At the end of the jacking solution step (Figure 2-32), the anchor-set step is conducted. Within 

this solution step, the ties from the ends of the “nubs” are removed, the tendon “nub” elements 

are removed, and new ties (in the deformed position) are created between the new tendon-ends 

and the center node of the concrete buttress. The final configuration at the end of anchoring is 

shown in Figure 2-33. This procedure is quite analogous to what occurs during construction. 

The tendon “nub” is the part of the tendon that is pulled out beyond the face of the concrete and 

essentially no longer exists for purposes of the completed structure (and in fact in most 

prestressing applications, it is simply cut off).
2-38

Figure 2-33 Deformed shape of hoop tendon system after “anchor set step”

Figure 2-34 and Figure 2-35 show the Hoop and Vertical Tendon Layout. Every tendon was 

modeled in Model 3.

ABAQUS has a built in PRESTRESS HOLD function, but this function is only for use in static 

analyses. It is also meant to keep the stress in some or all of the rebar constant. This is not for 

use in applying prestressing to tendons. ABAQUS also has a “PRE-TENSION SECTION” 

function, but this function merely associates a node with a section, and allows the user to define 

the normal components of the section. This can be used to define a pre-load to a section, but 

can be difficult to implement and set up future steps with tie-off of the tendons. Standard finite 

element practice consists of analysts defining temperature loads on sections causing a 

displacement of that section which is then calibrated to match the jacking displacement and 

allow the correct stress distribution to be modeled in the tendons. Both methods can be used, 

but because of the nature of the analyses to be completed, it was determined that using a 

temperature load would be the easiest to implement and control. 
2-39

Figure 2-34 Hoop tendon layout
2-40

Figure 2-35 Vertical tendon layout

A detailed, design oriented calculation prior to construction of the 1:4 Scale PCCV Model of 

tendon initial stress versus azimuth (including angular friction, wobble friction, and seating loss) 

was performed, with results shown in Table 2-2 and Figure 2-36. This was the target stress 

distribution used in the analysis [NUREG/CR-6810 2003].
2-41

Table 2-2 Tendon stress distribution for standard tendon behavior analysis (includes 

seating losses and assumed linearly varying with azimuth in between 

points.)

Azimuth Force (Newton)

85 334,625 (75,227 lb.)

45 381,526 (85,770 lb.)

360 323,648 (72,759 lb.)

270 230,512 (51,821 lb.)

180 323,648 (72,759 lb.)

135 381,526 (85,771 lb.)

95 334,625 (75,227 lb.)

85 334,292 (75,152 lb.)

Figure 2-36 Calculated tendon stress profile with losses for two tendons at given 

azimuth (blue and purple line)

0° 180°

270°

90°

F/W

M/S

A/L

E/H

80 90 100 110 120 130 140 150 160 170 180
2-42

2.4.1 Results and Comparisons from Prestressing Loads and Pressure Loads

Deformed shapes of the full model are shown in Figure 2-37 for after prestress and tendon 

anchorage step, and Figure 2-38 at pressurization of 3.6 times the design pressure or 3.6 Pd, 

which is incipient failure of the vessel.

A series of “plan-view” slice deformed shapes are shown in Figure 2-39 through Figure 2-46, at 

a model elevation of 4.68 meters (15.35 ft.) (roughly the mid-height of the cylinder). Based on 

comparisons to the test data these shapes and the magnitude of the displacements are in 

reasonably good agreement with observations from the limit state test (LST) and structural 

failure mode test (SFMT). 

A global plot of Maximum Principal Strains is shown for the concrete mid-thickness of the vessel 

in Figure 2-47. Studying the liner strain plots for pressures of 3.0 Pd, 3.3 Pd, 3.4 Pd, and 3.6 Pd, 

and comparing to known behaviors from the 1:4 Scale PCCV LST and SFMT, it can be 

concluded that many similar liner strain “hot spots” exist in the analysis as were observed in the 

test. For example, “hot spots” were observed: (a) near 0-degrees azimuth, cylinder midheight, 

(b) on either side of the equipment hatch (E/H) embossment, (c) on either side of the other 

penetrations (air lock (A/L), main steam (M/S), and feed water (F/W) penetrations).

Figure 2-48 and Figure 2-49 show hoop tendon stresses and Figure 2-50 and Figure 2-51 show 

vertical tendon stresses. The stress distribution plots show that the friction modeling strategy for 

Model 3 is very effective: the stress distributions after jacking and after anchorage are in 

agreement with design expectations and with observations and measurements from the test. 
2-43

Figure 2-37 Deformed shape after tendon anchorage (deformation scale × 500)
2-44

Figure 2-38 Deformed shape at 3.6 × Pd (deformation scale × 20)
2-45

Figure 2-39 Deformed shape at anchoring at elev. 4.68 m (15’-4 1/16”) (x500)

Figure 2-40 Deformed shape at Pd at elev. 4.68 m (15’-4 1/16”) (x500)
2-46

Figure 2-41 Deformed shape at 1.5 x Pd at elev. 4.68 m (15’-4 1/16”) (x250)

Figure 2-42 Deformed shape at 2.0 x Pd at elev. 4.68 m (15’-4 1/16”) (x100)
2-47

Figure 2-43 Deformed shape at 2.5 x Pd at elev. 4.68 m (15’-4 1/16”) (x50)

Figure 2-44 Deformed shape at 3.0 x Pd at elev. 4.68 m (15’-4 1/16”) (x30)
2-48

Figure 2-45 Deformed shape at 3.4 x Pd at elev. 4.68 m (15’-4 1/16”) (x30)

Figure 2-46 Deformed shape at 3.6 x Pd at elev. 4.68 m (15’-4 1/16”) (x30)
2-49

Figure 2-47 Maximum principal membrane strain in concrete at 3.6 x Pd
2-50

Figure 2-48 Stress in hoop tendons anchored at 90° after jacking before anchorage
2-51

Figure 2-49 Stress in hoop tendons anchored at 90° after anchorage
2-52

Figure 2-50 Stress in vertical tendons after jacking before anchorage (contours in psi)
2-53

Figure 2-51 Stress in vertical tendons after anchorage (contours in psi)

It should be noted that Figure 2-52 through Figure 2-60 plot Tendons H35, H53, and H68 which 

are hoop tendons located at elevations 4.572 m (15 ft.) , 6.579 m (21.58 ft.), and 8.280 m 

(27.165 ft.) on the PCCV model cylinder. The data from the test model are actually data from 

two different tests, as discussed in [NUREG/CR-6810 2003]. The Limit State Test (LST) 

occurred first and resulted in significant leakage of the test model such that further 

pressurization could not be obtained. The Structural Failure Mode Test (SFMT) was conducted 

second and included additional measures to ensure a leak-tight structure. The SFMT ended in 

structural failure of the vessel attributed to tendon failure. 
2-54

Figure 2-52 Abaqus analysis – hoop tendon H35 force 

Figure 2-53 LST test – hoop tendon H35 force

F/W

M/S

0° 180°

270°

90°

Tendon H35

Elevation: 4.572m 

Anchor: 270°

270°

A/L

E/H

200000

250000

300000

350000

400000

450000

500000

550000

600000

Force (newtons)

Azimuth (degrees)

(Load Cells and Average of Wire Strain Gages)

Design @

Max Tension

Design After

Seating

@ Max

Tension

After Seating

0 x Pd

1 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

2.6 x Pd

2 x Pd

0.9 x Pd

0 x Pd

270° 360, 0° 90° 180° 270°

F/W

M/S 0° 180°

270°

90°

Tendon H35

Elevation: 4.572m 

Anchor: 270°

270°

A/L

E/H
2-55

Figure 2-54 SFMT test – hoop tendon H35 force

Figure 2-55 Abaqus analysis – hoop tendon H53 force

0

100,000

200,000

300,000

400,000

500,000

600,000

Force (newtons)

Azimuth (degrees)

(Load Cells and Average of Wire Strain Gages)

Design @

Max Tension

Design After

Seating

@ Max

Tension

After Seating

0 x Pd

1 x Pd

2 x Pd

3 x Pd

3.5 x Pd

3.6 x Pd

3.5 x Pd

270° 360, 0° 90° 180° 270°

F/W

M/S

0° 180°

270°

90°

Tendon H35

Elevation: 4.572m 

Anchor: 270°

270°

A/L

E/H

0° 180°

270°

90°

Tendon H53

Elevation: 6.579m 

Anchor: 270°

270°

F/W

M/S

A/L

E/H
2-56

Figure 2-56 LST test – hoop tendon H53 force

Figure 2-57 SFMT test – hoop tendon H53 force

200000

250000

300000

350000

400000

450000

500000

550000

600000

Force (newtons)

Azimuth (degrees)

(Load Cells and Average of Wire Strain Gages)

Design @

Max Tension

Design After

Seating

@ Max

Tension

After Seating

0 x Pd

1 x Pd

2 x Pd

3 x Pd

3.2 x Pd

3.3 x Pd

2.6 x Pd

2 x Pd

0.9 x Pd

0 x Pd

270° 360, 0° 90° 180° 270°

0° 180°

270°

90°

Tendon H53

Elevation: 6.579m 

Anchor: 270°

270°

F/W

M/S

A/L

E/H

0

100,000

200,000

300,000

400,000

500,000

600,000

700,000

800,000

Force (newtons)

Azimuth (degrees)

(Load Cells and Average of Wire Strain Gages)

Design @

Max Tension

Design After

Seating

@ Max

Tension

After Seating

0 x Pd

1 x Pd

2 x Pd

3 x Pd

3.5 x Pd

3.6 x Pd

3.6 x Pd

3.5 x Pd

270° 360, 0° 90° 180° 270°

0° 180°

270°

90°

Tendon H53

Elevation: 6.579m 

Anchor: 270°

270°

F/W

M/S

A/L

E/H
2-57

Figure 2-58 Abaqus analysis – hoop tendon H68 force

Figure 2-59 LST test – hoop tendon H68 force

0° 180°

270°

90°

Tendon H68

Elevation: 8.280m 

Anchor: 90°

90°

F/W

M/S

A/L

E/H

200000

250000

300000

350000

400000

450000

500000

550000

600000

Force (newtons)

Azimuth (degrees)

(Load Cells and Average of Wire Strain Gages)

Design @

Max Tension

Design After

Seating

@ Max

Tension

After Seating

0 x Pd

1 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

2.6 x Pd

2 x Pd

0.9 x Pd

0 x Pd

90° 180° 270° 360, 0°

*Only one surviving strain gage at this azimuth.

*

90°

0° 180°

270°

90°

Tendon H68

Elevation: 8.280m 

Anchor: 90°

90°

F/W

M/S

A/L

E/H
2-58

Figure 2-60 SFMT test – hoop tendon H68 force

Figure 2-61 compares the Model 3 results to those from the LST [NUREG/CR-6810 2003] for 

the radial displacement at the equipment hatch. Figure 2-61 shows Model 3 radial displacement 

versus pressure compared to the LST measurement – the response prediction is in close 

agreement with the test.

0

100,000

200,000

300,000

400,000

500,000

600,000

700,000

Force (newtons)

Azimuth (degrees)

(Load Cells and Average of Wire Strain Gages)

Design @ Max

Tension

Design After

Seating

@ Max Tension

After Seating

0 x Pd

1 x Pd

2 x Pd

3 x Pd

3.5 x Pd

3.6 x Pd

3.5 x Pd

90° 180° 270°

*

360, 0° 90°

0° 180°

270°

90°

Tendon H68

Elevation: 8.280m 

Anchor: 90°

90°

F/W

M/S

A/L

E/H
2-59

Figure 2-61 Comparison of model 3 (ungrouted) result to LST test measurement at 

standard output location 14 which is the radial displacement at the center 

of the equipment hatch. Pressure is MPa with the grid divisions 

representing multiples of the design pressure (0.39 MPa). 

2.4.2 Grouted Tendon Model Method and Comparisons of Results

As previously described, for the ungrouted case, the tendons are attached to the concrete 

vessel with SLOT connector elements. These elements restrain motion between two nodes in 

two translational directions and allow motion in one direction (the rotations are unaffected). This 

one direction is assigned to be tangent to the curve at that point. In the motion direction, friction 

is assigned and is active in the jacking and anchor-set solution steps. 

In actual construction, all post-tensioned systems start out with ungrouted tendons. As such, the 

ungrouted tendon FEA modeling technique is pertinent to these systems in either case. For the 

grouted tendon FEA, after the tendons have been jacked and anchored, the SLOT connector 

elements are then replaced with BEAM connector elements, which restrain all six degrees of 

freedoms of the tendon node to the vessel-concrete node. Once these BEAM connector

elements are activated, the tendon can no longer move relative to the vessel, thereby simulating 

a grouted tendon. Note that it is restrained (“grouted”) in its deformed position relative to the 

concrete, and stays that way throughout the pressure load analysis. This method of simulating 

grouted tendons likely overestimates the bond between the tendon and the grout since the 

beam elements provide a rigid connection to the concrete nodes. Investigating the behavior of 

grouted tendons in containment vessels with particular emphasis on grout to tendon bond 

should be a topic of future research. 

Figure 2-62 and Figure 2-63 demonstrate how the SLOT elements are active during tensioning, 

which are then replaced with BEAM connector elements. 
2-60

Figure 2-62 During jacking – slot connector element connecting tendons to vessel

Figure 2-63 After anchorage– slot connector element removed and replaced with beam 

connector elements

To demonstrate the differences between grouted and ungrouted tendon behavior, comparisons 

are made in the following illustrations:

• Figure 2-64 Comparison of Grouted vs. Ungrouted Tendon Model 3 Results for Radial 

Displacements at Elev. 4.68 meters (15.35ft) (near Cylinder Mid-Height) 

• Figure 2-65 Comparison of Grouted vs. Ungrouted Tendon Model 3 Results for 

Tendon H35 Strains 

• Figure 2-66 Comparison of Grouted vs. Ungrouted Tendon Model 3 Results for 

Tendon H53 Strains 
2-61

• Figure 2-67 Comparison of Grouted vs. Ungrouted Tendon Model 3 Results for 

Tendon H68 Strains 

• Figure 2-68 Comparison of Grouted vs. Ungrouted Tendon Model 3 Results for 

Tendon V37 Strains 

• Figure 2-69 Comparison of Grouted vs. Ungrouted Tendon Model 3 Results for 

Tendon V46 Strains

Figure 2-64 Comparison of grouted vs. ungrouted tendon model 3 results for radial 

displacements at elev. 4.68 meters (near cylinder mid-height)

Ungrouted

Grouted
2-62

Figure 2-65 Comparison of grouted vs. ungrouted tendon model 3 results for tendon 

H35 strains

0.000

0.002

0.004

0.006

0.008

0.010

0.012

0.014

0.016

-90 -60 -30 0 30 60 90 120 150 180 210 240 270 300 330

Strains

Azimuth (degrees)

H35 Strains - Ungrouted

0xPd 1.1xPd 1.5xPd 2xPd 2.5xPd

3xPd 3.3xPd 3.4xPd 3.6xPd

F/W

M/S 0° 180°

270°

90°

Tendon H35

Elevation: 4.572m 

Anchor: 270°

270°

A/L

E/H

0.000

0.002

0.004

0.006

0.008

0.010

0.012

0.014

0.016

-90 -60 -30 0 30 60 90 120 150 180 210 240 270 300 330

Strains

Azimuth (degrees)

H35 Strains - Grouted

0xPd 1xPd 1.5xPd 2xPd 2.5xPd

3xPd 3.3xPd 3.4xPd 3.6xPd
2-63

Figure 2-66 Comparison of grouted vs. ungrouted tendon model 3 results for tendon 

H53 strains

H53 Strains Ungrouted 

0° 180°

270°

90°

Tendon H53

Elevation: 6.579m 

Anchor: 270°

270°

F/W

M/S

A/L

E/H

H53 Strains Grouted
2-64

Figure 2-67 Comparison of grouted vs. ungrouted tendon model 3 results for tendon 

H68 strains 

H53 Strains Ungrouted 

0° 180°

270°

90°

Tendon H53

Elevation: 6.579m 

Anchor: 270°

270°

F/W

M/S

A/L

E/H

H68 Strains Grouted
2

-65

Figure 2-68 Comparison of (a) grouted and (b) ungrouted results from model 3 for tendon 

V37, Azimuth 240°

0° 180°

270°

90°

Tendon V37Azimuth: 240°

270°

F/W

M/S

A/L

E/H
2

-66

Figure 2-69 Comparison of (a) grouted and (b) ungrouted results from model 3 for tendon V46, 

Azimuth 135°

0° 180°

270°

90°

Tendon V46Azimuth: 135°

90°

F/W

M/S

A/L

E/H
2-67

The radial displacement comparisons show that at the azimuths with largest radial

displacements (azimuths which will tend to drive the failure of the vessel) the grouted tendon 

solution is consistently larger than the ungrouted, though not by a large margin. At 3.6 Pd

pressure, the difference in radial displacement is approximately 7% (see Figure 2-64). The 

differences in strains in the tendons are a similar in magnitude. 

Two important general observations can be made with regard to the radial expansion of the 

cylinder and the hoop tendon behavior:

(1). Cylinder deformations are larger with grouted tendons, all other things being equal

(2). At high pressures, strains (and forces) are less well distributed along the tendons, and 

maximum strains are larger with grouted tendons versus ungrouted tendons

Both of these conclusions point to the broader conclusion that grouted tendon PCCVs will reach 

a failure limit state (either liner tear/leak or tendon rupture) at a lower pressure than ungrouted, 

all other things being equal. The vertical tendon behavior comparisons (Figure 2-68 and Figure 

2-69), in some ways, are even more illustrative of the differences between grouted and 

ungrouted tendons. At low pressures, as expected, the tendon strain distributions start out the 

same. By pressures of 2.5Pd and beyond, very noticeable differences occur. The differences 

correspond to the local distribution of vertical strain (including some bending effects near the 

wall base and near the springline) in the cylinder wall. Each increment of additional vertical 

strain in the cylinder wall adds to the strain in the adjacent tendon element one-for-one for the 

grouted case, but NOT for the ungrouted case. For the ungrouted case, the vertical tendons 

only respond to the total strain from top to bottom of the model, and along the cylinder, maintain 

uniform strain distributions. By 3.6Pd, this results in approximately 8% larger maximum strains 

in the grouted tendons versus the ungrouted ones.

For the vertical tendons (shown in Figure 2-68 and Figure 2-69) included during the grouted 

versus ungrouted tendon behavior discussion), the general stress levels and effects of friction 

show similar trends, but there are some differences in stress distribution observed in the dome, 

between the analysis and the test measurements, especially for Tendon V37. For Tendon V46, 

the trends between analysis and test compare well, including in the dome. Overall, it appears 

that the trends match between analyses and tests. If there were more results from the test along 

the vertical tendon, the match is expected to be better.

2.5 Conclusions

From Model 1, a model including a horizontal slice through the containment and only two 

tendons, it has been observed that the maximum tendon strains tend to be located at near 

where strain is maximum after prestress anchor set. This maximum is a function of the tendon 

to duct friction and the diameter of containment cylinder. For greater internal pressure levels, it 

appears that the stress and strain maximum moves in response to containment dilation and 

tendon yield. The circumferential slip of tendons relative to the concrete (after jacking and 

anchor-set, during pressurization) is about 3 millimeters (0.118 in.) at 3.4 x Pd. If this were a fullscale PCCV under the same design conditions, the slip would most likely be greater than 12

millimeters (0.472 in.). Ideally there would be a direct scaling. However, due to the smaller 

radius of the 1:4 Scale Test as compared to a full scale containment the angular friction of the 

scale model is 30% to 50% larger than it would be in a prototype. As such, it is possible that the 

non-scalable angular friction would cause the full-scale slip to be a bit higher than 4x the slip 

seen in the 1:4 analyses. The use of contact surfaces to model tendon friction works well when
2-68

the number of finite elements in the model is small, and the relevant friction data are readily 

available. There appears to be significant computational expense associated with this method 

and other techniques are better for larger models.

Accordingly, Model 3, a global shell analysis model of the entire 1:4 Scale PCCV, employs a 

novel SLOT-connector based approach for modeling the interaction between the tendons and 

the reinforced concrete structure. Tendon stress distribution results show that this friction 

modeling strategy is effective since stress distributions after jacking and after anchorage are in 

reasonable agreement with design and with observations from the test [NUREG/CR-6810 

2003]. Hoop tendon strain contours at 3.6Pd indicate that the largest tendon strains occur near 

the anchorages, in the 0-6 degree midheight zone of the cylinder, and in the 135-degree zone of 

the cylinder (see Figure 2-65 and Figure 2-66). The 1:4 scale test model failed in this 0-6 degree 

range so the high tendon strains in this region for the simulation are expected. The Ungrouted 

Model 3 analyses are in reasonable agreement with the LST and SFMT. 

From both the Model 1 and Model 3 studies, plus literature review, several conclusions have 

been reached about comparing “ungrouted vs. grouted” behavior, from the FE Modeling 

perspective. First, modeling only the “force effects” of the tendons (no stiffness, no separate 

elements), is only appropriate for elastic analysis (and is equally appropriate for ‘greased vs. 

grouted’). It is inappropriate for severe accident analysis. Second, modeling ‘greased’ tendons 

by imbedding tendon elements in the concrete (strain compatibility) can produce approximate 

results, but can lead to premature prediction of tendon rupture, because tendon strains increase 

one-to-one with vessel wall strains. Instead, this strain would re-distribute a small amount thus

lowering the maximum value (e.g. Figure 2-22). Furthermore, it is important to ensure the initial 

tendon stress distribution is realistic and accurate, and this can be laborious. Conversely, 

modeling grouted tendons as ‘bonded’ elements (say rebar subelements) will produce realistic 

results for most studies, but it is essential that the proper stress distribution is implemented in 

the tendons prior to grouting; and generally speaking this is a laborious manual procedure. One 

such way to do this would be to assign stress to individual sections of the tendon by assigning 

local strains calculated based on the estimated stress state of the tendon, and then iterating to

ensure that with equilibrium, the stress distribution along the tendon was distributed properly.

Additionally, as mentioned above, the method of algebraically tying the tendon nodes to the 

concrete causes the tendon strain to accumulate in a localized manner that would likely not be 

the case in a real system as the grout began to fail. Finally, it is far more computationally 

‘elegant’ to model the frictional interaction between the tendons and the ducts using

friction/contact-surfaces or the “SLOT”-type connectors to reach the prestress distribution 

automatically. This approach makes it a better predictor of true behavior during pressurization 

for ungrouted tendons. Additionally, the SLOT connector approach allows for a straightforward

‘grouting step’ for grouted tendons, though the shortcomings of not capturing grout failure 

mentioned above also apply here. 

Approximating the bond between the grout and the tendon with rigid BEAM elements likely 

overestimates the bond between the tendon and the grout. The error would likely arise from the 

lack of simulated bond failure between the grout and the tendon, which would tend to occur at 

higher internal pressure levels. Despite this potential shortcoming, useful conclusions can still 

be obtained, yet more research on detailed grouted tendon behavior is warranted. In general, 

for PCCVs, comparing grouted to ungrouted tendon behavior the maximum cylinder 

deformations are larger with grouted tendons, all other things being equal. Also at high 

pressures, strains (and forces) are less well distributed along the tendons for grouted tendons 

this is because local increments of strain in the vessel wall must track one-for-one with tendon 

strain increments; not so for ungrouted tendons. Accordingly, maximum strains are larger with 

grouted tendons versus ungrouted tendons and this will likely lead to containment failure at 

lower pressure levels with grouted versus ungrouted prestressing systems.
3-1

3 COMPARISON OF POST-TENSIONING AND INSERVICE 

INSPECTION OF GROUTED AND UNGROUTED SYSTEM

An industry-wide review of post-tensioning methods, standards, and inspection for application to 

PCCVs has been performed. This section describes the installation of post-tensioning systems, 

and the differences between grouted and ungrouted systems from the perspective of 

construction.

Prestressed concrete (which includes both pre- and post-tensioning strategies) involves the use 

of high strength steel tensioned against the concrete [Collins and Mitchell 1997]. The tensioning 

results in a self-equilibrating system of internal stresses which improves the response of the 

concrete to external loads. For example, if a plain concrete member (of say 34.47 MPa (5,000 

psi) compressive strength) were subjected to axial tension, the concrete would crack when the 

average tensile stress reaches about 2.068 MPa (300 psi), and this failure occurs at relatively 

small deformation (tensile strains of approximately 0.0001). As described in [Collins and Mitchell 

1997], if standard rebar is added (for example 1.5%, by area, of 413.69 MPa (60 ksi) steel), the 

member still cracks at nearly 2.068MPa (300 psi) average stress, and at nearly the same strain, 

but the member can now continue to resist loads until the reinforcement yields. If instead, the 

member contains about 41.53 kg/m3 (70 lb/cy) of rebar and 20.76 kg/m3

(35 lb/cy) of highstrength prestressing steel (about 0.25% by area), the member now sustains 5.171 MPa (750 

psi) average stress without cracking (2.5 times that of the reinforced member), and also has 

substantial ductility. This is demonstrated by the following calculations which use ACI design 

guidelines [ACI 2011]. 

For a reinforced concrete member, suppose the concrete compressive strength, fc’, is equal to 

34.47 MPa (5,000 psi). Young’s Modulus, Ec, is calculated using equation below: 

(3-1)

The cracking strength, fcr, of concrete is also calculated using fc’:

(3-2)

Assuming a rebar reinforcing ratio, ρs, of 1.5%, a Young’s Modulus of Steel, Es, and gross area 

of the section, Ag, the average member stress, fmember, at member cracking is calculated below:

(3-3)

For a prestressed or post-tensioned concrete member, it is assumed that the conventional rebar 

reinforcing ratio, ρs, is 0.5%, and that the post-tensioning percentage, ρp, is 0.25%. A typical 

prestressing stress, fp, is assumed:

(3-4)

' 6 4700 27579 (4 10 ) E f MPa x psi c c = =

' 4 1.944 (282 ) cr c f f MPa psi ≈ =

1.11 2.158 (313 ) cr g

member cr

sgs

g

c

f A f f MPa psi A E

A

E

ρ

= =× =     +

 

f MPa ksi p =1.241 (180 )
3-2

The prestressing force, Fp, is calculated using the equation below:

(3-5)

The compressive force, fcp, in the concrete is then calculated:

(3-6)

Using the above, the contributions to member tensile strength, Ts, include the recovery of the 

concrete pre-compression plus the original concrete tensile strength plus the steel contribution. 

The average member stress at member cracking is therefore:

(3-7)

(3-8)

(3-9)

Ts is equal to the tensile strain, εs, from applied load times Young’s Modulus times steel area, 

As. εs is approximately equal to the total change in concrete stress divided by Ec.

(3-10)

(3-11)

Substituting these results back into the equation for the average member stress at member 

cracking:

(3-12)

Therefore, prestressing can be highly effective for crack prevention, and for generally producing 

structures which are strong, tough, and stiff. 

Since prestressing of concrete structures prior to service loading is very effective for minimizing 

or eliminating cracking at service loads (and to control deflections), it tends to also result in 

more slender, lighter structures. All of these attributes can be beneficial to structure design. For 

example, a prestressed one-way floor slab can have a span-to-depth ratio of about 24 to 1, 

which is about 60% more than the ratio possible with a non-prestressed one-way slab. And for a 

given span, the amount of concrete in the prestressed slab will be less than two-thirds that of 

the reinforced concrete slab. More than 50% of bridges are now constructed of prestressed 

concrete. In North America, there are over 500 plants that produce precast, pre-tensioned 

structural elements. The repetitive nature of the production process—together with the 

450 F Af A p pgp= = × ρ g

3.10 ( 450 ) p

cp

g

F

f MPa psi A

− = =− −

282 pgp s

member

g g

A f T f A A

ρ

= ++

450 282 s

member

g

T f A

=++

6

450 282 0.00018

4 10

cp cr

s

c

f f

E

ε

− + +

≅ ==

×

( ) 6 0.00018 30 10 40 T EA s sss = = ×× × + = ε ρ ρ p s g g A A

fmember = + += 450 282 40 5.323 MPa psi (772 )
3-3

controlled environment of the plant—results in elements with high-quality concrete and 

dimensional control.

Because parking structures tend to be subjected to very corrosive environments, the use of

high-quality concrete—prestressed to control cracking—has become standard practice. 

Approximately 75% of parking structures are built with prestressed concrete. Corrosive 

environmental conditions also exist for pier/waterfront structures, so similar design approaches 

are often employed.

In 1936, Freyssinet demonstrated that prestressed concrete cylindrical structures can resist 

considerable internal pressure without cracking or leakage [Freyssinet 1936]. This principle has 

been put into practice for a wide variety of prestressed concrete pressure vessels. In particular, 

prestressed concrete (inclusive of all pre-loading techniques) is the construction method of 

choice for more than one third (37 out of 104) of the reactor containment vessels in the United 

States, and an even higher percentage worldwide. While in the United States, these vessels 

generally have a thin steel liner to ensure leak-tightness, in some countries, these vessels are 

unlined, yet still have remarkable pressure retaining capacity without cracking or leakage.

3.1 Comparison of Post-Tensioning Systems

3.1.1 Post-Tensioning Methodology

Concrete post-tensioning systems in buildings typically consist of ungrouted single-strand 

tendons used in floor and roof slabs. Post-tensioning is also used in beams and in some cases 

columns but post-tensioned frame members are not as commonly used as slabs, especially in 

seismically active areas. The choice between ungrouted tendons and grouted tendons is 

principally a function of the relative ease of inspection and repair or replacement potential in the 

nuclear industry. Both systems are allowed by the governing criteria. The design codes 

governing prestressing design for buildings are provided at the end of this chapter. 

Tendons typically consist of single seven wire strands in a high density polyethylene sheath that 

is extruded onto the strands. Prior to extruding the sheath the strands are typically coated with a 

grease or wax containing corrosion inhibiting admixtures. A typical ungrouted single strand 

tendon is shown in Figure 3-1.
3-4

Figure 3-1 Typical ungrouted single strand tendon

Prestressing in waterfront structures is typically pre-tensioned rather than post-tensioned, so the 

installation procedures are not very relevant to PCCVs, however the corrosion protection and 

monitoring needs do have some common issues (as discussed later).

Bridges were among the first uses of prestressing, and today more than 50% of bridges are now 

constructed of prestressed concrete – some using pre-tensioning (with precast elements 

manufactured at casting yards), and some using post-tensioning. California and Florida, in 

particular, rely heavily on cast-in-place post-tensioned elements, and of course these 

construction methods are the most analogous to PCCV construction. Nearly all post-tensioned 

applications for bridges are grouted, but there are notable exceptions called “external 

prestressing”. Generally, external prestressing does not mean external to the bridge – it means 

external to the solid concrete (i.e., not embedded). External prestressing is sometimes run 

inside of concrete box-girders and is harped (to apply moments counteracting vertical loads) 

through the use of deviator blocks and saddles. 

Provided in Figure 3-2 is a typical post-tensioning detail for a bridge or other large civil structure 

(not floor slabs as described previously for buildings). Also provided on the drawing are call-outs 

for typical ASTM material standards. For comparison, a representative section of a PCCV with 

rebar and post-tensioning detail is shown in Figure 3-3. Additionally, a picture of the 1:4 Scale 

PCCV model under construction, with rebar and tendon ducts visible, is shown in Figure 3-4 . In 

both the bridge and PCCV detailed sections, it is apparent that the steel in the sections are 

dense, requiring careful concrete installation to ensure no voids are present. 

Dead End Anchorage

Monostrand Intermediate Anchorage

Stressing Anchorage

Typical monostrand post-tensioning system components (VSL post-tensioning).

7-wire strand with corrosion 

preventive coating completely filling 

annular space Anchorage

Plastic Cap

Mortar Plug

Grease Fitting

Plastic tube with watertight 

connection at anchorage

Tight plastic sheath
3-5

Figure 3-2 Typical post-tensioning detail for bridges or other civil structures
3-6

Figure 3-3 Typical post-tensioning detail for PCCV from [NUREG/CR-6810 2003] for (a) 

Vertical section (b) Horizontal section
3-7

Figure 3-4 PCCV wall mock-up from [NUREG/CR-6810 2003]

Post-tensioning ducts are placed along with rebar cages, and tied into place. Most often these 

ducts are metallic, but plastic ducts are also deemed acceptable by some jurisdictions (PCCV 

ducts are currently required to be metallic). After concrete is poured and cured, tendons are 

jacked and anchored. Methods vary as to what measures are taken for corrosion protection. 

Tendons are always provided with a layer of grease, and if tendons are to be grouted, this may 

be the only protective measure needed. Bridge and PCCV construction specifications limit the 

amount of time which can occur between tendon placement and grouting. These limits are 

generally on the order of days, and preferably, less than 4 weeks. ASME Boiler and Pressure 

Vessel Rule CC-4282 [ASME 2010a] requires that the grout be placed no more than 7 days 

after tensioning if no measures are taken to control the corrosion on the surface of the tendons 

and no more than 30 days if corrosion control measures are taken. If the possibility exists for 

longer time duration between tendons existing loose in the ducts, and grouting, then additional 

measures, such as oil injection into the ducts are followed. A detailed comparison of duct 

practice is contained in subsection 3.1.4.
3-8

3.1.2 Key Difference between Post-Tensioning Grouted and Ungrouted Tendon 

Systems

Post-tensioning systems are usually one of several proprietary systems. Some of the most 

frequently used are the VSL multi-strand system [VSL 2001], the Freyssinet K-Range System

[Freyssinet 2010], the Dywidag post-tensioning bar system [Dwyidag 2008], and the BBR 

system [BBR 2013]. Additionally, the EPR design uses Freyssinet C-Range Strands [Freyssinet 

2010]. The most commonly used system in PCCVs is a multi-strand system. All of the 

aforementioned systems utilize multiple strands with nominal diameters between 24 mm and 

12.7 mm (0.9 inches and 0.5 inches). Each tendon can have up to 55 strands depending on the 

system, but the most common number used in PCCVs is seven. The EPR design uses a 55 

strand C-Range system.

In grouted post-tensioned construction the ducts are grouted as soon as possible after 

stressing. The objective is to fill the duct completely with material that provides an alkaline 

environment for corrosion protection of the prestressing steel and to provide a continuous bond 

between the tendon strand and the duct. One of the functions of the duct is to transfer bond 

between the grout within the duct and the concrete surrounding the duct. To minimize this timeframe, the tendons are generally inserted into the ducts just prior to stressing. Grout usually 

consists of a mixture of cement and water (water/cement ratio of about 0.50) together with a 

water-reducing admixture and an expansive agent (for PCCV construction, the w/c ratio must be 

less than 0.45 by weight [Regulatory Guide 1.107 2011]. For larger diameter ducts (in PCCVs), 

grout may also contain fillers such as sand, fly ash, or pozzolans. (There are strict requirements 

for grout materials for PCCVs to restrict, components of the grout may not damage or interact 

with the prestressing tendon [Regulatory Guide 1.107 2011].) Satisfactory grout should 

maintain fluidity during the grouting, exhibit minimum bleeding and segregation, should not 

shrink, should have adequate strength, and should not contain detrimental amounts of 

chlorides, nitrates, sulfides, or other corrosion inducing compounds. 

Grout is injected at the low points of the tendon or at the ends of the member. Venting tubes are 

provided at the high points of the tendon. If ducts are not properly vented, pockets of air may be 

trapped at high points of the duct. Freezing of water that may collect in these air pockets can 

result in serious deterioration of the structure and may lead to corrosion. The long-term 

durability of a grouted, post-tensioned structure depends on the success of the grout material 

and grouting operation. Specifications for this highly specialized and critical procedure are given 

by the Post-Tensioning Institute [PTI 2012]. At present, Finland’s STUK (Radiation and Nuclear 

Safety Authority), and France’s EDF (Électricité de France) perform tests on full size and partial 

mock-ups to ensure grouted tendon construction methodology is adequate. 

Ungrouted post-tensioned construction is nearly identical to the aforementioned procedure for 

grouted construction except that more care and stricter specifications are applied to the coatings 

of the strands to prevent corrosion. It should be noted that the tendon strand coatings (or the 

material injected into ducts, such as grease or oil) can have some influence on the friction loss 

characteristics of the tendons. The use of emulsible oils to reduce friction losses and provide 

temporary corrosion protection for tendons prior to grouting is acceptable practice. Research 

shows that if the tendon is stressed while the oil is still fresh, the coefficient of friction can be 

reduced by as much as 15%. In addition, it has been shown through bond tests that the strength 

of specimens with oiled tendons is similar to or better than unoiled tendons for specific types of 

oils [Lüthi, Diephuis et al. 2008]. However, care must be taken in selecting the oil used, because 

some oils can significantly damage the bond between the strand and the grout [Davis, Tran et 

al. 1993]. Flushing the ducts with water is not guaranteed to remove enough of the oil to prevent 

damaging the bond between the strand and the grout. In addition, flushing is not allowed in 

PCCV construction. 
3-9

Differences in friction are generally addressed during design by the engineer using tendon 

system data provided by the prestressing manufacturer. Other than this potential difference in 

friction, the installation procedures, and the stress distribution conditions for grouted and 

ungrouted tendons are the same when a new structure is put into service. Additional differences 

may evolve over time as the structure ages. Differences between PCCV requirements and 

those in the PCI [PCI 2010] are discussed in Section 3.1.3.

If stress in a post-tensioned concrete structure (stress on the concrete) is held constant for 

some time, the strain increases, and this phenomena is referred to as creep. If the strain is held 

constant for some time, the stress will decrease. This is referred to as stress relaxation. Both 

types of time dependent deformation are commonly referred to as creep, though the time 

dependent deformation resulting from prestressing is a stress relaxation phenomenon. Creep is 

generally estimated using creep tests of a specific concrete mix, but even with such tests, creep 

results in the large, poured structure can be rather variable compared to those of test 

specimens under laboratory conditions. Another aging phenomenon that affects the stress state 

in a PCCV over time is shrinkage, which also strongly depends on the composition of the 

concrete – the total amount of water in the mix being especially important. The quality of the 

aggregate is also important with hard, dense, stiff aggregates of low absorption (e.g., hard 

limestone or granite), resulting in less shrinkage. Other time-dependent effects in concrete are 

the increase in stiffness and strength of concrete over time, sometimes changing by between 

5% and 30% between 1-year and 40-years. 

One of the main effects generally considered in assessing time-dependent behavior of posttensioned structures is relaxation of prestressing steel. It is generally held within the design 

industry (for non-nuclear applications) that tendon relaxation is negligibly small (2%-3%) if the 

initial stress applied to the steel is less than 0.55fy, and that the little relaxation that does occur, 

occurs within the first few hours after stressing [Collins and Mitchell 1997]. However, as tendon 

stress levels rise (Figure 3-5) relative to fy, relaxation effects can become much more significant, 

and can have a substantial time-varying component (10% or more as shown in the figure). US 

Regulatory Guide (RG) 1.35.1 recommends the initial stress be kept below 70% of the 

guaranteed ultimate tensile strength [Regulatory Guide 1.35.1 1990].
3-10

Figure 3-5 Summary of potential tendon relaxation demonstrating the decrease in 

relaxation time for increasing applied stress [Structural Research Series 

No. 237 1962].

All of these time-dependent effects tend to create accumulating additional “losses” in the 

amount of effective stress in the tendons, and in turn pre-compression of the concrete and the 

effects will occur in both grouted and ungrouted tendon systems; however, with the 

3-dimensional complexity of PCCVs, they may occur with different stress distributions in grouted

versus ungrouted PCCVs. Creep and relaxation effects will be largest in zones where the 

effective prestress is largest. In ungrouted tendon PCCVs such longer term losses will tend to 

be more smoothly distributed spatially over the structure than in grouted tendon PCCVs. Similar 

trends were seen in Section 2 of the report when the PCCV was subjected to overpressurization, and the stresses and strains were more smoothly distributed over the structure. 

For grouted tendon PCCVs, development of long term losses will occur locally in the zones 

where they are most prevalent and be less likely to spread to zones which started with lower 

effective prestress. We have not examined the potential effects of this (due to scope 

constraints), but it does represent a worthwhile analysis modeling topic. 

3.1.3 Industry and Structural Code Comparison of Post-Tensioning 

Methodologies

The governing codes for designing and installing post-tensioning systems vary between 

structure type across the industry. 

General structures (buildings and other non-nuclear, non-bridge structures) are designed to the 

International Building Code [IBC 2012] (IBC, similar to the long-standing Uniform Building 

Code), and this takes many cues from the ACI-318 Code [ACI 2011]. Bridge structures are 

designed to State and Local standards, which generally have their roots in the ACI-318 Code, 

but are augmented by local experience and local practice.
3-11

Nuclear containment designs and construction are governed by the ASME Boiler and Pressure 

Vessel code, Section III, Division 2 [ASME 2010a] and the Inservice Inspection governed by the 

ASME Boiler and Pressure Vessel code Section XI [ASME 2010b]. It should be noted that the 

ASME code Section XI references the ACI-349 Code [ACI 2013] and other NRC regulatory 

documents for certain issues. 

For installation and corrosion-inhibiting specifications, there is some commonality between 

industries; this mostly stems from the fact that much of the best testing, construction method, 

and corrosion-performance data comes from manufacturers of the tendon systems. For 

example, VSL provides the industry with a significant database of specifications and best 

practices (see for example [VSL Report Series 5 2002]) and similar information is available from 

Freyssinet [Freyssinet 2010]. 

One document which compares concrete (reinforced, R/C, and prestressed, P/C, concrete) 

design code practice is the National Cooperative for Highway Research Program’s “NCHRP 

Report 549 – Simplified Shear Design of Structural Concrete Members” [NCHRP 2006]. 

Although the focus of the report is on shear design, it provides a comprehensive summary of 

various international design codes used for prestressed concrete design. The relevant codes 

are:

• ACI 318 (2002)

• AASHTO STD

• AASHTO LRFD

• CSA (both 1994 and 2004, in which many provisions changed)

• EC2 (1991 and 2003)

• DIN (2001)

• JSCE

• AASHTO for Segmental Bridges

• Frosch (2002)

3.1.4 Post-Tensioning Duct and Grouting Practice

A comparison of post-tensioning duct practice and grouting requirements as related to 

construction practice for various industries is contained herein. Specifically, ACI 318 [ACI 2011], 

Caltrans [Standard 2010], AASHTO [AASHTO 2013], FHWA [FHWA 2013], ASME [ASME

2010a], and US Reg. Guide 1.107 [Regulatory Guide 1.107 2011] are discussed, and then 

some comparison tables, where useful, are provided. A more detailed discussion of grout 

chemistry and requirements is contained in Section 4. 

3.1.4.1 ACI 318 Requirements

For standard building structures, ACI 318, Sections 18.16 and 18.17 [ACI 2011] have the 

following general requirements for ducts (for grouted tendons) and sheathing (for ungrouted

tendons):

• Ungrouted tendons:

o A sheath is required to surround the prestressing steel

o The sheath must be filled with a material to inhibit corrosion
3-12

o The sheath is required to be watertight and continuous over the entire length of 

the tendon

o When in corrosive environments, the sheathing has to be connected to all 

stressing, intermediate, and fixed anchorages in a watertight fashion

• Grouted tendons:

o The ducts must be mortar-tight and nonreactive with concrete, prestressing steel, 

grout, and corrosion inhibitor

o The inside diameter of the duct must be at least 6.35 mm (1:4 in.) larger than the 

prestressing steel diameter for single wire, single strand, or single bar tendons

o For multiple wire, multiple strand, or multiple bar tendons, the inside crosssectional area of the duct must be at least two times the cross-sectional area of 

the prestressing steel

o The ducts must be maintained free of ponded water if the members to be grouted 

are exposed to temperatures below freezing prior to grouting.

Grouting requirements are defined in Section 18.18, and are as follows:

• Grout can consist of Portland cement, sand, and water.

• Portland cement shall meet ACI standards.

• Water shall meet ACI standards.

• Sand will meet Standard Specification for Aggregate for Masonry Mortar with gradation 

variation modified as necessary for workability.

• Admixtures can have no negative effect on grout, steel, or concrete, and calcium 

chloride is not allowed.

• Proportions for grout shall be based on either results of tests on fresh and hardened 

grout, or on prior documented experience with similar materials and equipment under 

comparable field conditions.

• Water content shall be minimum necessary for pumping, and shall not exceed 0.45 by 

weight.

• Water cannot be added to increase grout flowability that has been decreased by delayed 

use of the grout.

• Grout must be mixed in equipment that can provide continuous mechanical mixing and 

agitation that will produce uniform distribution of materials, passed through screens, and 

pumped in a manner that will completely fill the ducts.

• Temperature of members at time of grouting needs to be above 1.67 o

C (35 oF) and held 

above 1.67 o

C (35 oF) until field cured 5.08cm (2 inch) cubes of grout reach a minimum 

compressive strength of 5.52 MPa (800 psi).
3-13

• Grout temperatures shall not exceed 32.22 oC (90 oF) during mixing and pumping.

3.1.4.2 Caltrans Requirements

The duct requirements from Caltrans in the Standard Specifications from the State of California 

Department of Transportation [Standard 2010] can be found in Section 50, Prestressing 

Concrete.

As Per 50-1.02D and 50-1.03A(3), ducts must:

• Be galvanized rigid ferrous metal.

• Be fabricated with either welded or interlocked seams.

• Be mortar tight.

• Have sufficient strength to maintain their correct alignment during placing of concrete.

• Have positive metallic connections at joints between sections that do not result in angle 

changes at the joints.

• Have waterproof tape at the connections.

• Have bends that are not crimped or flattened.

• Have ferrous metal or polyolefin transition couplings connecting the ducts to anchorage 

system components.

• For multi-strand tendons, ducts must have an internal area of at least 2.0 times the net 

area of the tendon. When pull-through method is used, internal area must be at least 2.5 

times the net area of the tendon.

• For PT bars, ducts must have an inside diameter at least 9.525 mm (3/8”) larger than the 

diameter of the bar.

• Have an outside diameter not exceeding 50 percent of the girder web width.

• Ducts must be injection grouted after prestressing. Grout must consist of cement and 

water and may contain an admixture if authorized, in accordance with Section 50-1.02C, 

“Grout.”

• Be accurately placed.

• Be securely fastened to prevent movement during concrete placement.

• After installation, duct ends must be covered to prevent water or debris from entering.

For ducts with a vertical profile change of six inches or more, the ducts must be vented. The 

vents must be placed within six feet of every high point in the duct profile. In addition, as per 50-

102E, the vents must:

• Be at least 12.7 mm (1⁄2 in.) diameter standard pipe or suitable plastic pipe.
3-14

• Be connected to the ducts using metallic or plastic structural fasteners. Plastic 

components must not react with the concrete or enhance corrosion of the prestressing 

steel and must be free from water soluble chlorides.

• Be mortar tight and taped as necessary.

• Provide a means for injection of grout through the vents and for sealing the vents.

3.1.4.3 AASHTO Requirements

The AASHTO requirements for ducts can be found in Section 5.4.6, Ducts, of Chapter 5, 

Concrete [AASHTO 2013]. The general requirements for ducts are listed in Article 5.4.6.1 and 

the corresponding commentary, which states:

• Ducts for tendons shall be rigid or semi-rigid either galvanized ferrous metal or 

polyethylene, or they shall be formed in the concrete with removable cores.

• The radius of curvature of tendon ducts shall not be less than 20.0 ft., except in the 

anchorage areas where 3.66 m (12.0 ft.) may be permitted.

• Where polyethylene ducts are used and the tendons are to be grouted, the bonding 

characteristics of polyethylene ducts to the concrete and the grout should be 

investigated.

• The effects of grouting pressure on the ducts and the surrounding concrete shall be 

investigated.

• The maximum support interval for the ducts during construction shall be indicated in the 

contract documents and shall conform to Article 10.4.1.1 of the AASHTO LRFD Bridge 

Construction Specifications.

• The use of polyethylene duct is generally recommended in corrosive environments. 

Pertinent requirements for ducts can be found in Article 10.8.2 in AASHTO LRFD Bridge 

Construction Specifications.

• Polyethylene duct should not be used on radii under 9.144 m (30.0 ft.) because of its 

lower resistance to abrasion during pulling-through and stressing tendons.

The size requirements for the ducts are listed in Article 5.4.6.2, and states:

• The inside diameter of the ducts shall be at least 6.35 mm (0.25 in.) larger than the 

nominal diameter of single bar or strand tendons.

• For multiple bar or strand tendons, the inside cross-sectional area of the duct shall be at 

least 2.0 times the net area of the prestressing steel.

• Where tendons are to be placed by the pull-through method, the duct area shall be at 

least 2.5 times the net area of the prestressing steel.

• The size of ducts shall not exceed 0.4 times the least gross concrete thickness at the 

duct.
3-15

• The pull-through method of tendon placement is usually employed by contractors where 

tendons exceed 121.92 m (400 ft.) in length.

3.1.4.4 FHWA Requirements

The US Federal Highway Administration defines requirements for ducts in post-tensioning 

systems in the Post-Tensioning Tendon Installation and Grouting Manual [FHWA 2013]. Ducts 

are described in Section 2.3, Ducts, of Chapter 2, Post-Tensioning System Materials and 

Components. Ducts are available in different materials for different applications and types of 

tendons. Originally, duct was considered primarily as a means of forming a void through the 

concrete for the tendon and little attention was paid to the possible role of the duct as a barrier 

to corrosive agents. Largely as a consequence of finding voids in grouted tendons, more 

emphasis is now placed on the quality, integrity and continuity of the duct as a corrosion barrier 

in itself. This has resulted in a move toward the use of high density plastic ducts in some states. 

Nevertheless, previous duct materials are still available and their use continues in other regions. 

Consequently, the following recommendations should be adapted as appropriate to meet local 

needs and conditions.

The duct size is specified for strand tendons in Section 2.3.1.1 of the FHWA Post-Tensioning 

Tendon Installation and Grouting Manual, and for bar tendons in Section 2.3.1.2. Specifically:

• For strand tendons, the internal area of circular ducts should be at least 2.25 times the 

net area of the strands.

• When pull through method is used, internal duct area should be at least 2.50 times the 

net area of the strands.

• In case of space limitations, the minimum duct area may be 2.00 times the strand area 

for relatively short tendons up to approximately 30 m (100 ft.) long.

• Oval “flat” ducts are commonly used for transverse tendons comprising up to 4 strands 

of 15 mm (0.6 in.) diameter in deck slabs of box girders. The internal clear dimensions of 

oval duct should be a minimum of 25 mm (1 in.) vertically and 76 mm (3 in.) horizontally.

• For tendons containing a single post-tensioning bar the internal duct diameter should be 

at least 6.35 mm (1:4 in.) greater than the maximum outside dimension of the bar. A 

greater clearance may be preferred or be necessary for some applications. Examples of 

this use would be to provide greater tolerance for temporary bars or to accommodate 

bridges with slightly curved alignments.

Requirements are also given for various types of ducts, as summarized below. Corrugated steel 

ducts are specified in Section 2.3.2.1, smooth rigid steel pipe in section 2.3.2.2, corrugated 

plastic ducts in section 2.3.2.3, and smooth high density polyethylene pipe for external tendons 

in Section 2.3.2.4.

• Corrugated steel ducts

o Ducts are spirally wound to the necessary diameter from strip steel with a 

minimum wall thickness of 0.45 mm (26-gauge) for ducts less than 66 mm (2-

5/8”) diameter or 0.6 mm (24-gauge) for ducts of greater diameter.
3-16

o The strip steel should be galvanized to ASTM A653 with a coating weight of G90. 

(It is worth noting that the Farley NPP had issues with galvanized materials 

causing problems. See Generic Issue 118 [NUREG-0933 2010])

o Ducts should be manufactured with welded or interlocking seams with sufficient 

rigidity to maintain the correct profile between supports during concrete 

placement (Figure 3-6 (a)). 

o Ducts should also be able to flex without crimping or flattening.

o Joints between sections of duct and between ducts and anchor components 

should be made with positive, metallic connections that provide a smooth interior 

alignment with no lips or abrupt angle changes.

• Smooth rigid steel pipe

o Smooth steel pipes should conform to ASTM A53 "Standard Specification for 

Pipe, Steel, Black and Hot-Dipped, Zinc Coated, Welded and Seamless", Grade 

B Schedule 40.

o When required for curved tendon alignments (e.g. deviation saddles and similar) 

the pipe should be pre-fabricated to the required radius (Figure 3-6(b)).

• Corrugated plastic ducts

o Corrugated plastic duct (Figure 3-7) to be completely embedded in concrete 

should be constructed from either polyethylene or polypropylene.

o The minimum acceptable radius of curvature should be established by the duct 

supplier according to standard test methods.

o The duct should have a minimum wall thickness of 2.0 mm + 0.25 mm (0.079 in. 

+ 0.010 in.). 

o Ducts should have a white coating on the outside or should be of white material 

with ultraviolet stabilizers added.

• Smooth high density polyethylene pipe for External Tendons

o HDPE smooth pipe is available in different diameters, wall thickness, physical 

and chemical properties. There is significant variability in commonly available 

materials. It is very important that it has satisfactory properties for handling, 

storage, installation and durability for the application. The color is normally black 

from a small amount of carbon in the material, to protect against degradation 

from ultraviolet light. The wall thickness, diameter and physical strength 

(Hydrostatic Design Basis) should be sufficient to initially withstand grouting 

pressures. In the long term it should not deteriorate or split. The requirements 

should be in accordance with AASHTO LRFD Bridge Construction Specifications.
3-17

Figure 3-6 Spiral Wound Steel Duct and Rigid Steel Pipe

Figure 3-7 Corrugated plastic duct
3-18

The requirements for plastic fittings and connections for internal tendons are defined in Section 

2.3.2.5, which state:

• All duct splices, joints and connections to anchorages should be made with couplings 

and connectors that produce a smooth interior duct alignment with no lips or kinks.

• Special duct connectors may be used in match-cast joints between precast segments 

and similar situations if necessary to create a continuous, air and water-tight seal. 

• Duct tape should not be used to join or repair ducts or make connections.

• All fittings and connections between lengths of plastic duct and between ducts and steel 

components (e.g. anchors or steel pipe) should be made of materials compatible with 

corrugated plastic ducts. 

• Plastic materials should contain antioxidant stabilizers and have an environmental stress 

cracking of not less than 192 hours as determined by ASTM D1693 "Standard Test 

Method for Environmental Stress-Cracking of Ethylene Plastics", Condition C.

External tendon duct connection requirements are listed in Section 2.3.2.6 and are as follows:

• Connections between sections of plastic pipe should be made using heat welding 

techniques or with mechanical couplers per the manufacturer's recommendations or as 

approved by the Engineer. Connections should have a minimum pressure rating of 1.0 

MPa (145 psi) and produce a smooth interior alignment with no lips or kinks.

• Connections between external HDPE pipe and steel pipe embedded in the concrete 

should be made using circular sleeve (boot) made of Ethylene Propylene Deine

Monomer (EPDM) having a minimum (working) pressure rating of 1.0 MPa (145 psi). 

EPDM should have 100% quality retention as defined by ASTM D1171 "Standard Test 

Method for Rubber Deterioration-Surface Ozone Cracking Outdoors or Chamber 

(Triangular Specimens)" Ozone Chamber Exposure Method B. The minimum wall 

thickness should be 10mm (3/8 inch) reinforced with a minimum of four ply polyester 

reinforcement. Sleeves should be secured with 10 mm (3/8 in.) wide power seated, 316 

stainless steel band clamps, using one on each end of the sleeve (boot) to seal against 

leaking grout. The power seating force should be between 356 N and 534 N (80 lbf and 

120 lbf). Alternatively, connections may be made using mechanical couplers with plastic 

components made of approved plastic resins meeting the same requirements as for 

external plastic pipes and metal components of grade 316 stainless-steel. Mechanical 

connections should meet the same pressure rating requirements (above) and have seals 

to prevent grout leaks.

• Steel and plastic pipe may be connected directly when the outside diameters do not vary 

by more than + 2 mm (0.08 in.). A reducer or spacer should be used when outside this 

tolerance. When installed correctly, a single band clamp around each end of the sleeve 

should be sufficient. Double banding may be necessary to fix an apparent leak of air, 

water or grout.

The requirements for shrink sleeves are listed in section 2.3.2.7:

• In some cases, external tendon connections may be enhanced by the use of shrink 

sleeve wrap overlaying the connection and portions of adjacent plastic and steel pipes.
3-19

• Shrink sleeves may be used in aggressive environments where connections may be 

directly exposed to de-icing salts or other contaminants entering through expansion 

joints or other similar openings.

• Shrink sleeves should consist of an irradiated and cross linked, high density 

polyethylene backing with an adhesive layer that will withstand 66° C (150° F). Sleeve 

materials should meet the requirements listed in Table 3-1:

Table 3-1 Physical properties required for shrink sleeves.

Property Test method

Minimum requirements

Internal 

application

External 

application

Units

Fully recovered 

Thickness* 2.3 (92) 2.8 (111) mm (mils)

Peel strength: ASTM D1000 5.0 (29) 8.0 (46) KN/M (lb per inch)

Softening Point: ASTM E28 72 (162) 102 (216) °C (°F)

Lap Shear: DIN 30 672M 60 (87) 40 (58) MPa (psi)

Tensile Strength: ASTM D638 20 (2,900) 24 (3,480) MPa (psi)

Hardness: ASTM D2240 46 52 Shore D

Water Absorption: ASTM D570 <0.05% <0.05%

*Note: The fully recovered thickness is the thickness after installation using heat.

The requirements for shipping, handling, and storage of ducts are listed in Subsection 2.3.3, 

and are as follows:

• Duct made from galvanized strip steel may be prefabricated or fabricated on site as 

necessary. 

• Plastic duct may be shipped in coils or in bundles of straight lengths.

• In order to avoid inadvertent introduction of contaminants or debris, it is recommended 

that the ends of duct coils or bundles be protected and covered during shipping and 

storage. 

• Special temporary end caps may be used to seal the ends of individual ducts. 

• Plastic ducts should be protected from sunlight, ultraviolet degradation, crushing and 

excessive bending until installed in the bridge. 

• All ducts and pipes should be stored in a dry location, on a raised platform, protected 

from weather and contamination.

The acceptance criteria for duct materials are defined in Section 2.3.4, and are as follows:

• All duct materials (metal or plastic) ducts should comply with the requirements of 

AASHTO LRFD Bridge Construction Specifications or the project specifications.
3-20

• In general, post-tensioning duct will be acceptable if it meets the requirements of 

"Acceptance Standards for Post-Tensioning Systems", Section 5, "Sheathing", PTI, 1998 

and for corrugated plastic ducts, FIB Bulletin #7,"Corrugated Plastic Ducts for Internal 

Bonded Post-Tensioning Systems", Article 4.2 "System Approval Testing" Stage 1 and 

Stage 2.

• Key features for acceptance (according to PTI) for internal tendons are:

o Duct cast into concrete should withstand at least 3.0 m (10ft) of concrete fluid 

pressure.

o Duct shall not dent more than 3 mm (1/8 in.) under a concentrated force of 0.45 

kN (100 lbf) applied using a 13 mm (0.51 in.) diameter (#4) reinforcing bar.

o Where prestressing steel is pre-installed in the duct, the duct shall withstand at 

least 1.5 m (5 ft.) of concrete fluid pressure and resistance to denting is not 

required.

o Duct with a diameter greater than 50mm (2 in.) shall not deflect more than 75 

mm (3 in.) when a 6 m (20 ft.) length is supported at its ends, although where 

tight radii are required, more flexible duct may be permitted.

o Plastic duct should withstand the above at 38°C (100°F) except that longitudinal 

stiffness requirements may be reduced by 50% if the installation support spacing 

is reduced 50% from that for steel duct.

o The above do not apply to ducts stiffened with bars, mandrels or inflatable tubes.

• For acceptance, it is recommended that three successful and successive tests for each 

type of duct should comply with the above requirements.

• FIB Bulletin #7 [FIB-7 2000] sets out procedures for the approval of a corrugated plastic 

duct system on the basis of a series of stages, using the same assembly of the system. 

The stages are:

o Practicability of Assembly - the actual assembly of the ducts in a rebar cage

o Water Tightness - by an air-pressure test of the same assembly prior to 

concreting

o Stressing/Friction - by jacking and releasing at one end then the other

o Grouting Test and Wear of Duct - no voids nor significant wear after grouting 

(autopsy)

o Electrical Resistance Test - not less than 1 kilo-Ohm resistance between the 

reinforcing and post-tensioning steel (for electrically isolated systems - prior to 

autopsy).

• FIB Bulletin #7 [FIB-7 2000] is suitable for qualifying new systems that have not been 

used before. The tests are not meant as project specific tests or production tests but as 

system approval tests and therefore, need to be carried out only once for certification or 

approval of the system.
3-21

• Key acceptance requirements for external tendons:

o Duct (HDPE pipe) including all welds, splices, grout fittings and connections 

should be vapor tight and capable of withstanding 1 MPa (145 psi) grout 

pressure.

o For verification, an assembly containing plastic and steel pipe and connections 

may be pressure tested as follows: (1) Condition the assembly by sustaining an 

internal pressure of 1MPa (145 psi) for 3 hours. (2) After conditioning, the 

assembly should retain an internal pressure of 1MPa (145 psi) for five minutes 

with no more than 0.1 MPa (15 psi) reduction.

• It is recommended that a system supplier provide full documentation including:

o Technical documents and drawings of general assembly of the system and 

details of components

o Instructions and method statements for installation, stressing and grouting

o A quality assurance plan covering production, shipping, handling, storage and 

installation of the system

o Instructions for surveillance and maintenance of the system in service

• For acceptance and approval of a post-tensioning system, all components tests and 

results of post-tensioning system approval tests should be carried out by an independent 

approved body or testing laboratory. This testing should be completed prior to 

submission of Shop Drawings and other related documents to the Engineer for approval.

• Not all the above tests are standardized and are not formal requirements. The above 

may be used for guidance. Requirements for individual projects should be considered on 

a case-by-case basis. Proposals should be prepared by the Contractor for the approval 

of the Engineer.

• On site, the Contractor should maintain a complete record of all documentation, test 

reports, shipping dockets and approvals. Copies should be provided to the Inspector 

(CEI) to ensure compliance. Also, it is recommended that for multi-strand tendons, 

internal tendon ducts be checked using a torpedo prior to installing tendons (see 

Chapter 3).

3.1.4.5 ASME Requirements

The ASME [ASME 2010a] Requirements (Section III, Division 2, Code for Concrete 

Containments) contains requirements for ducts in post-tensioned systems. Section CC-2440, 

Nonload-Carrying and Accessory Materials, contains section CC-2441, Tendon Ducts, 

Channels, Trumpets, and Transition Cones. Important points from this section are as follows:

• The material shall be ferrous and shall not cause harmful electrolytic reactions with the 

prestressing element.

• Ducts and channels shall be strong enough to retain their shape and resist irreparable 

damage during vessel construction.
3-22

• Ducts and their joints shall prevent intrusion of cement paste from the concrete.

• When used for single element tendons, the inside diameter of ducts shall be at least 6 

mm (1:4 in.) larger than the nominal diameter of the tendon.

• When used for multiple element tendons, the inside cross-sectional area of ducts shall 

be at least twice the nominal net area of the enclosed tendon.

• When filler injection is required, ducts shall have openings at both ends. Vents shall be 

provided at high points. Drains shall be provided at low points.

• Ducts and duct joints shall be capable of withstanding an internal pressure greater than 

the initial pumping pressure when adequately backed by concrete.

Section CC-4452 lists requirements for tendon installation related to tendon ducts and channels, 

and Section CC-5423 lists requirements for pre-placement and post-placement of the ducts. Of 

note:

• The Construction Specification shall specify the tolerances for position and alignment. 

• Tendon ducts and channels shall be adequately supported against displacement during 

concreting. 

• Open ducts shall be protected by capping or plugging to prevent entry of concrete or 

other deleterious material. 

• All joints shall be made tight against the inleakage of mortar or appreciable water from 

the fresh concrete. 

• The Construction Specification shall specify the temporary corrosion protection system, 

if any, and the construction procedures shall define the method for its application. 

• All water and debris shall be removed from ducts prior to installation of tendons. 

• Tendon ducts shall be examined to ensure compliance with requirements of the 

Construction Specification as to type, diameter, and wall thickness. 

• The frequency of examination shall be specified in the Construction Specification. 

• Tendon ducts shall be examined for position and alignment in accordance with CC-4452. 

• After installation, but prior to concrete placement, the ducts shall be visually examined 

for damage, including holes and cracks, and for dents and ovalling which may affect 

required minimum clear aperture. The criteria for the visual examination shall be given in 

the Construction Specification. 

• Prior to and following concrete placement, all ducts shall be examined to ensure that 

minimum clear aperture is provided.

Grouting requirements are listed in section CC-4280, and are as follows:

• Equipment for measuring and batching grout must be accurate within plus or minus 3%.
3-23

• Equipment for mixing must produce grout of uniform consistency, free from lumps and 

undispersed cement.

• Grout must be continuously agitated until pumped.

• Pump must be capable of exerting the required maximum pressure. Safety device must 

prevent pump from exerting a discharge pressure over 2 MPa (300 psi). 

• The pump shall not draw in air with the grout.

• A screen shall be used to prevent lumps in the grout.

• Equipment must prevent leakage and segregation of grout.

• Grouting shall be carried out as promptly as possible after tensioning. 

• The total time of exposure of the tendon steel to other than a controlled environment 

prior to grouting shall be specified in the Construction Specification and shall not exceed 

30 days, nor shall the period between tensioning and grouting exceed 7 days unless 

special precautions are taken to protect the tendon against corrosion. 

• The methods or products used in that case shall not jeopardize either the effectiveness 

of the grout as a corrosion inhibitor or the bond between the prestressing element and 

the grout. 

• The Designer shall evaluate the potential corrosive effects of the local environmental

conditions to determine if additional restrictions on the above limits must be established 

in the Construction Specification.

• Flushing prior to grouting is not recommended.

• Maximum time after mixing within which the grout can be used is established by tests 

that ensure the intended reaction of the admixtures continues throughout injection, and 

the time is two hours less than that required for the initial set of the grout.

• The temperature at any point in the tendon duct during grouting shall be above 4.4o

C 

(40o

F). 

• The end anchorages and the tendon duct shall be maintained above 4.4o

C (40oF) for a 

period of 48 hr after grouting or until the grout reaches a minimum of 5.52 MPa (800 psi) 

compressive strength. 

• The grout temperature shall not exceed 37.8o

C (100oF) during mixing and pumping 

• All openings, air vents, and drains shall be hermetically sealed after grouting to prevent 

the ingress of water or any corrosive agents. 

• When an expansive grout is used, all high point openings shall remain open to permit 

free expansion until grout sets.
3-24

3.1.4.6 US RG 1.107 Requirements for Grouted Tendons

The following is a list of key requirements from the US RG 1.107 [Regulatory Guide 1.107 

2011]:

• The grout must completely fill the duct on hardening to prevent water from collecting and 

freezing.

• Preassembled tendon-sheathing assemblies are to be protected before concrete against 

a corrosive environment during assembly handling, storage, transportation, placement, 

and tensioning.

• Prior to tendon placement in the duct, it must be ascertained that the duct is free of 

obstructions, moisture, and other deleterious substances.

• Ducts must be mortar-tight, and nonreactive with concrete, prestressing steel, and grout.

• Minimize the time duration between tensioning and grouting processes.

• The duct needs to be sized such that the tendons can be inserted and tensioned without 

undue difficulty. 

• All the vents and drains should be checked for obstructions prior to grouting, and the 

ducts should be maintained free of water. 

• The duct / sheath may only be made of metal currently.

• The pH of the grout at the inlet and the outlet of the duct should be maintained above 

11.6 (or 12 if the allowable chloride content is 199.8 mg/L (200 ppm))

• The water/cement ratio of the grout must be under 0.45 by weight.

• Grouting equipment should include a mixer that is capable of continuous mixing and can 

produce a grout free of lumps and undispersed cement. Applicants should conduct tests 

to demonstrate the optimum range of mixing time.

• A screen with clear openings smaller than 3 mm (0.12 in.) for standard grout and 5 mm 

(0.2 in.) for grout with thixotropic additive should be placed between the mixed grout and 

the pump. If lumps of cement remain on the screen, the batch should be rejected.

• Temperature at any point in the duct during grouting should be above 4.4o

C (40oF). 

• End anchorages and tendon duct shall be maintained above 4.4o

C (40oF) for a period of 

48 hours after grouting or until grout reaches a minimum compressive strength of 5.52 

MPa (800 psi).

• Grout temperature should not exceed 32.2o

C (90oF) during mixing and pumping 

operations unless the applicant can establish through testing that a higher temperature 

will not adversely affect the quality and performance of the grout.
3-25

3.1.4.7 Comparison Table

Table 3-2 is a comparison of some of the key parameters from the above sections compared 

directly.

Table 3-2 Comparison of key parameters for ducts and grouting.

Code Duct material 

allowed Size of duct

Time duration 

between 

tensioning and 

grouting

W/C ratio 

for grout

Minimum and 

maximum 

temperatures 

during grouting

ACI Metal

Inside crosssectional area 

at least two 

times the crosssectional area 

of the 

prestressing 

steel

N/A Under 

0.45

Above 1.67o

C 

(35o

F), Under 

32.2o

C (90oF)

AASHTO

Ferrous 

Metal, 

Polyethylene, 

concrete 

formed by 

removing 

cores

Inside crosssectional area 

at least 2x the 

net area of the 

prestressing 

steel

Minimize Under 

0.45 N/A

Caltrans

Galvanized 

Rigid Ferrous 

Metal

Inside crosssectional area

at least 2x the 

net area of the 

prestressing 

steel

N/A

Less than 

5 Gallons 

of water 

per 94 

Pounds of 

Cement

Protect from 

quick setting and 

freezing

FHWA

Steel, 

Polyethylene, 

and 

polypropylene

Inside crosssectional area 

at least 2x or 

2.5x the net 

area of the 

prestressing 

steel

Minimize Under 

0.45 N/A

ASME Ferrous Metal

At least 2x the 

nominal area of 

the tendons

Minimize Under 0.53

Over 4.44o

C 

(40o

F), Under 

37.78o

C (100oF)

US Reg. 

Guide Metal

Allow for 

Insertion of 

Tendons

Minimize Under 

0.45

Over 4.44o

C 

(40o

F), Under 

32.2o

C (90oF)

 3 It should be noted that 2013 version of the ASME Code requires w/c less than 0.45.
3-26

3.1.5 Summary and Conclusions

Looking across the building industry, design methods for post-tensioning systems (both grouted 

and ungrouted) are similar, but differences can be observed at the detailed level of establishing 

installation and inspection specifications. These topics are discussed in more detail in the next 

section.

At the time that post-tensioning (tendon stressing) is completed, structures which are to remain 

ungrouted have nearly the same stress distribution and structural performance character as 

structures whose ducts will be grouted. An exception to this are systems where extra oil or 

grease are injected as this extra grease may slightly reduce angular and wobble friction 

coefficients of ungrouted systems relative to those which will be grouted. But after grouting, and 

with the onset of passage of time, differences in stress distributions will likely occur due to timedependent changes in prestress (and pre-compression of concrete) due to creep, shrinkage, 

strength-aging of concrete, and due to tendon steel relaxation. These changes will occur in 

similar magnitude between grouted and ungrouted systems, but may have different spatial 

distributions between the two types of structures because after grouting, grouted tendons have 

strain compatibility with concrete, while ungrouted tendons do not.

Differences in behavior caused by these spatial differences have not been investigated, but they 

are expected to be relatively small. The only substantial structural response performance 

differences occur at pressures large enough to overcome prestressing and crack the concrete 

walls of the containment. 

Despite possible corrosion protection advantages, polymeric tendon ducts (HDPE or similar), 

will require significant study before their use in the nuclear industry will be accepted. The ability 

of these non-metallic duct materials to withstand the large hydrostatic pressure generated by 

the massive concrete pours common in the nuclear industry, the radiation flux, and the elevated 

temperature conditions should all be verified before use, and research into answering these 

questions may be required. 

3.2 Comparison of Inservice Inspection Methods for Grouted and 

Ungrouted Tendon Systems

3.2.1 Needs of Post-Tensioned Monitoring

Since their first use 70-80 years ago, many advances in tendon system design and construction 

have been introduced to primarily aid structure performance, especially for poor or catastrophic, 

rare occasions. Although tendon corrosion issues have occurred infrequently in PCCVs over the 

years, and never has a corrosion related problem led to a structure failure in a PCCV, the data 

and experiences for bridges are much more extensive having a wider range of observed 

performance. This industry provides strong historical cases that support the need for tendon 

monitoring, and its objectives. After the discussion of historical cases from the bridge industry, a 

discussion of problems observed in the PCCV industry is provided. The focus is on observed 

issues with grouted and ungrouted post-tensioning systems in PCCVs around the world. 

3.2.1.1 Evidence Provided by Bridge Industry on Need for Monitoring

Since bridges are often found in even more corrosive environments than buildings (over water, 

and in marine environments), there have been decades of lessons learned about corrosion 

protection, inspection, and grouting methods. 
3-27

Historically, post-tensioned concrete bridges have been considered to need less maintenance 

than steel bridges, but after collapses of a few post-tensioned concrete bridges in Europe, it was 

realized that post-tensioned systems do have long-term risks. Many such bridges have 

exhibited corroded tendons and severe wire fractures caused by ingress of water and chloride 

ions into partially grouted ducts. Grouted tendons are difficult to inspect, typically only by using 

invasive inspection due to the lack of reliable nondestructive techniques. Invasive inspection 

requires excavation of concrete and only exposed tendons among the total number of tendons 

can be inspected. In the case of ungrouted tendons, the condition of the whole length of 

exposed tendons can be checked without exact information of the location of corroded tendons, 

but for grouted tendons, this is not the case.

Post-tensioned bridges can be susceptible to the ingress of moisture, particularly at beam and 

deck ends, high and low sections of tendon profiles and at construction or segment joints. The 

complete grouting of ducts is demanding and voids can be formed. This increases the risk of 

oxygen, moisture, salts and carbon dioxide entering the ducts and corrosion of the tendons may 

result. The corrosion may remain hidden until the resulting loss of strength is sufficient to 

promote either external evidence of the deterioration or a sudden failure. 

In a 1998 report on research on the “Durability of Precast Segmental Bridges” sponsored by 

AASHTO in co-operation with the United States Federal Highway Administration [NCHRP 

1998], it was noted that “... the corrosion problems with precast segmental and post-tensioned 

bridges in the United Kingdom were principally due to poor design choices and poor quality 

construction rather than a significant intrinsic susceptibility to corrosion. The study found no 

evidence of corrosion or other durability problems with precast segmental bridges in the United 

States”. The care taken in the design and construction of PCCVs prevents these issues from 

causing corrosion problems directly, but as discussed in Section 3.2.1.2, corrosion of tendons 

has been observed in PCCVs in the United States.

But other researchers have written that the last statement must be treated with caution as visual 

inspection alone will not necessarily provide any evidence of deterioration of post-tensioning 

systems. Investigations during 1999 and 2000 found severe corrosion and failed tendons in the 

Niles Channel Bridge, Mid-Bay Bridge and the Sunshine Skyway Bridge (Figure 3-8) in Florida. 

On January 12, 2001 the FHWA issued a memorandum stating: “Historically the corrosion of 

post-tensioned bridges in the United States has not appeared to be a significant problem. That 

viewpoint has recently been brought into question by the discovery of failed and corroded posttensioning tendons in the State of Florida”. The memorandum stated that a nationwide notice 

should be issued to “alert States to the post-tensioning corrosion problems” and “to recommend 

the expedited inspection of their highest risk post-tensioned structures”.

Voids in themselves may not pose a problem as long as all the prestressing wires are covered 

with grout. If wires are exposed then any ingressing water and salts will present a threat to longterm durability of any structure. The risk is related to how well voids are sealed. Significant 

corrosion of tendons has been found in investigations in the UK, France, Denmark, Germany, 

Italy and Japan.

United States grouting and tendon protection practice has been shaped, in large part, by tendon 

corrosion issues found in Florida bridges. In Florida, inspections of bridges revealed a 

significant number of voids at anchorages (such as shown in Figure 3-8) resulting from an 

accumulation of bleed water. In a large Florida interstate segmental interchange, a number of 

partially grouted tendons and ungrouted tendons were found. It has been suggested in 

[Research Report 0-1405-9 2004] that the tendon corrosion problems in Florida bridges resulted 

from:
3-28

• Voids associated with accumulation of bleed water at tendon anchorages

• Re-charge of ungrouted tendon anchorages with salt water or surface drainage during 

construction

• Leakage through end anchorage protection details

• Quality of grout installation and grout material

• Splitting of polythene ducts

• Deficiencies in implementation and inspection of grouting procedures

All of these issues, to varying degree, are generic to tendon post-tensioning systems, and thus 

are relevant to PCCVs. In addition, they have helped shape nuclear regulatory policy. RG 1.107 

requires that the duct be completely full of grout to prevent voids. There are strict standards for 

the grout material, and a quality installation must be assured. In addition, at present only metal 

ducts are allowed in PCCVs, and the implementation and inspection of grouting procedures 

must be qualified.

Figure 3-8 Corrosion observed near ends of tendons from the Sunshine Skyway 

Bridge

Following the announcement in September 1992 that the UK DoT would not commission 

bridges with grouted duct post-tensioning, the UK Concrete Society and the UK Concrete Bridge 

Development Group formed a Working Party to carry out research and prepare a report on what 

was needed to be done to improve the durability of post-tensioned concrete bridges. The first 

edition of Technical Report 47 “Durable Post-Tensioned Concrete Bridges (TR47) was 

published in 1996 and the second, updated edition 8, in 2002. A revision of this document was 

released as TR 72 in September of 2010 [Technical Report 47 2010]. The European Standards 

EN 445, EN 446, and EN 447 [EN 445 2008, EN 446 2008, EN 447 2008] now cover the basic 

requirements for grout, and TR 72 reviews test methods for grout materials and duct systems, 

includes a specification for duct and grouting systems, and references international and 

European standards where they exist. Some of the measures described in the reports have 

been considered by government agencies outside the UK, including Florida DOT, Virginia DOT 

and the Boston Central Artery/Tunnel Project, prior to these agencies issuing new/revised 

specifications. The first edition of TR47 recommended new standards and practices for the 

design and construction of durable grouted post-tensioned concrete bridges. It covered the key 

elements of design, detailing, materials, grouting and certification for installation. This resulted in 

1996 in the lifting of the moratorium on post-tensioned construction in the UK, other than for 

segmental construction. TR 72 includes recommendations for buildings as well as bridges.

In conclusion, nuclear industry policies for PCCV post-tensioning can benefit from observations 

and procedures for other structures. For bridges, corrosion problems with post-tensioning have 
3-29

not only been found in the United Kingdom, but also in Europe, Japan, the United States, and 

other countries. A great deal of attention has been paid by engineers to design and stressing 

operations, and perhaps, not enough attention has been paid to the subsequent grouting 

operations.

Isolated voids in and of themselves are not a major problem. Where the ducts are well sealed 

from the atmosphere, the presence of even a thin layer of grout over the surface of the tendons 

has provided effective corrosion protection. What is not known is for how long this layer of grout 

will continue to protect the tendons, especially where adjacent wires have no covering. 

Corrosion is caused by the ingress of moisture and other deleterious materials through 

vulnerable areas such as anchorages, segment joints or cracks. The difficulty of detecting 

corrosion, and the absence of visual evidence of deterioration in most cases, means that visual 

inspection alone is not sufficient to determine the condition of these structures. Further durability 

problems may occur with internally grouted tendons in the future as the corrosion protection 

provided to-date begins to break down and moisture and salts gain access over time. High risk 

structures will require careful management. The problems encountered demonstrate the need to 

improve detailing, and the specifications of materials, site operations, testing and protection of 

post-tensioning systems. 

3.2.1.2 Evidence Provided by PCCV Industry on Need for Monitoring

This section focuses on observed issues in PCCVs in the United States and around the world, 

focusing as much as possible on recent lessons learned with post-tensioning systems.

A potential serious issue in PCCVs, as in bridges mentioned above, is the corrosion of 

prestressing tendons. There have been instances of corrosion discovered in prestressed tendon 

systems within the United States. In October of 1999, the US NRC released Information Notice 

99-10, Revision 1: Degradation of Prestressing Tendon Systems in Prestressed Concrete 

Containments [NRC 1999]. During the 20th-year surveillance of the prestressing system of 

Calvert Cliffs Unit 1, a low lift-off value compared to the prestressing force for one of the three 

randomly selected vertical tendons. During the subsequent investigation, it appeared that 

strands were breaking during further testing. As a result 100% of the vertical tendons were 

tested. The licensee ended up having to replace 63 out of 202 vertical tendons in Unit 1 and 64 

of 204 vertical tendons in Unit 2. The corrosion in the tendons was attributed to uneven shim 

stack heights on the anchor-heads, spalling and cracking of concrete beneath the anchor-head 

base plates, free water in the bottom grease caps, poorly drained top anchorage ledges, and 

the absence of filler grease in various areas.

Another area of concern related to tendons which was documented in Information Notice 99-10, 

Revision 1, is the effect of high temperature loads causing excessive relaxation of tendons over 

time. Robert E. Ginna Nuclear Power Plant, Virgil C. Summer Nuclear Station, Turkey Point 

Station, have reported lower-than-predicted prestressing forces for vertical, hoop, and dome 

tendons. Investigations and analyses indicated that prestress losses ranged from 15.5% to 20% 

as compared to the initial design values which predicted tendon loss values in the range of 4% 

to 12%. These excessive relaxation values were due to the tendons in the containment being 

exposed to high temperatures (32.2o

C (90oF)) for extended periods of times. Many plants 

around the United States are typically exposed to average temperatures above 32.2o

C (90oF), 

but not all plants will necessarily have excessive loss of prestressing force because of 

conservative estimates in the design, frequent unsystematic re-tensioning of tendons, and 

improper use of a method of trending measured tendon forces, or any combination of the 

aforementioned reasons. This relaxation can be observed in ungrouted systems through direct 
3-30

testing, but this is not the case for grouted tendon systems, in which lift-off tests cannot be 

performed on the grouted tendons.

In February and March of 1985, the US NRC released Information Notice No. 85-10: 

Post-Tensioned Containment Tendon Anchor Head Failure, and its supplement 1, respectively

[1985]. Within this notice, the licensees were notified that Farley Unit 2, Bellefonte Units 1 and 

2, and Byron Units 1 and 2 had failure of anchor heads within their post-tensioned system, and 

water was discovered within the greased tendon systems. The cause of the failure of the anchor 

heads was deemed to be hydrogen-stress cracking, which in turn was caused by the sustained 

tensile loads in conjunction with the presence of atomic hydrogen. The hydrogen was produced 

as a result of corrosion. As a result of this, the licensees had to remove all of the vertical tendon 

caps, and test all those anchor heads with magnetic particle testing. In addition, where water 

was found in the hoop and dome tendons, testing had to be performed. At least 28 anchor 

heads had to be replaced [Naus, Oland et al. 1999, Braverman, Miller et al. 2004]. 

Another potential area of concern in the construction of PCCVs is the accurate construction of 

the containment vessel. It is difficult to ensure that in-situ construction will match the approved 

design of the structure. This concerns material quality, construction quality, and ensuring best 

practices in construction are followed. 

An example of a documented case where in-situ construction didn’t match design was observed 

at the EPR being constructed at the Flamanville site in the lower Normandy region of France. At 

this location, it was observed that there were positioning anomalies of prestressing sheaths for 

the reactor building inner containment wall. In addition, the layout of concrete reinforcement did 

not match the design [Nuclear 2010]. As a result, EDF had to demonstrate that the noncompliance was not detrimental to the performance of the structure. In addition, in the 

construction of subsequent lifts, ASN checked to ensure provisions were implemented to 

prevent the repeat of anomalies [Garnsey, Nickson et al. 2011]. One of the key 

recommendations from Garnsey, Joyce, and Nickson is the need to ensure that subcontractors 

are of high quality and are experienced in nuclear construction or are taught the necessary 

special skills related to quality, traceability, and documentation. Vilkangas of STUK also 

recommends the need for qualified engineers and contractors with the new construction of 

PCCVs [Nuclear 2010]. 

Finland has also had experience with having to change the location of prestressing tendons to 

account for changes in polar crane design. In order to do this, a new non-linear 3D analysis had 

to be completed to verify design margins cover local decreases in capacity [Nuclear 2010]. 

Both Finland and France have had extensive experience with the use of grouted tendons in their 

PCCVs. One of the standard practices which is beginning to be used, but not yet required by 

regulation, is the testing of partial or full mock-ups of the tendon grouting process [Debattista 

and Dubois 2011, Louhivirta 2011]. As mentioned previously, grout is injected at the low points 

of the tendon or at the ends of the member. Venting tubes are provided at the high points of the 

tendon. If ducts are not properly vented, pockets of air may be trapped at high points of the 

duct. Freezing of water that may collect in these air pockets can result in serious deterioration of

the structure and may lead to corrosion. The long-term durability of a grouted, post-tensioned 

structure depends on the success of the grouting operation. EDF considers injection techniques 

to be qualified if three successive samples of each tendon type comply with the required criteria

[Debattista and Dubois 2011]:
3-31

• Horizontal and vertical tendons shall be grouted in one phase, and gamma tendons shall 

be grouted in two phases, with the first phase for the vertical part of the tendon, and the 

second phase for the dome part of the tendon

• Ducts shall not contain free water nor shall the grout display potentially damaging cracks 

when the ducts are opened

• For purely vertical tendons and the vertical part of gamma tendons no voids shall be 

present

• For pure horizontal tendons, deviated horizontal tendons, and dome sections of gamma 

tendons, defects shall not have dimensions that expose strands.

• For the descending part of deviated tendons, the volume of voids in the descending part 

of the duct shall be limited.

Despite this testing, there is no guarantee that every tendon in the PCCV will be free of voids 

and free water. It is essential to ensure that corrosion is not developing over time within the 

prestressing system. However, this practice does help ensure quality construction practices, 

and is seen as current best-practice.

3.2.2 Different Industry and Code Requirements for Monitoring Loss of Prestress

This section describes different industry requirements for monitoring loss of prestress to 

highlight different approaches depending on industry needs, and to assess best practice for 

PCCVs. Industry practice for buildings and waterfront structures are summarized, and then a 

discussion of PCCV practice is provided. 

3.2.2.1 Building and Waterfront Structure Industry Practice

For conventional buildings (and waterfront structures) there are generally no long-term 

monitoring (strain-gaging of strands or load-cell measuring, etc.) measures in-place after a 

structure is placed into service. Tendon forces (generally referred to as Pjack) are measured and 

achieved during stressing, and records are provided of jacking operations, including Pjack and 

total elongation (length of the tendon pulled out of the end of the duct). This is the extent of 

monitoring performed for ungrouted tendon systems.

The same standard methods pertain to grouted systems, but then additional measures are 

taken during grouting. Temperature limits (on grouting construction days) are enforced, because 

it is not usually possible to grout ducts successfully if the temperature of the surrounding 

concrete is less than about 4.44o

C (40oF). Prior to grouting, the ducts are blown out using oilfree compressed air to remove debris from the ducts. Sometimes the ducts are flushed out with 

water prior to grouting (this is prohibited for containments). If blockage of a duct occurs during 

grouting, the duct needs to be flushed out immediately by injecting water into the closest air 

vent, against the direction of grouting. Records of grouting pressures, volume of grout used, 

temperatures, and other details are required and kept. High grouting pressures (e.g., over 1.5 

MPa, or 200 psi) may indicate blockage in the duct and may cause segregation of the grout or 

splitting of the concrete surrounding the duct. 

Generally, loss of prestress is not monitored during the service life of the structure, but problems 

can sometimes be suspected based on visual, tell-tale signs of loss of prestress, including 

excessive deflections over time, or local cracking on the concrete surface.
3-32

3.2.2.2 Nuclear Structure (PCCV) Practice

The two regulatory guides for the inspection of post tensioned tendons installed in concrete 

containment structures are:

• RG 1.35 Inservice Inspection of Ungrouted Tendons in Prestressed Concrete 

Containment Structures

• RG 1.90 Inservice Inspection of Prestressed Concrete Containment Structures with 

Grouted Tendons

Both regulatory guides provide guidance for meeting the requirements of General Design 

Criterion (GDC) 53, “Provisions for Containment Testing and Inspection” of 10 CFR 50, 

“General Design Criteria,” Appendix A, which require that the containment be designed to permit 

(1) periodic inspection of all important areas and (2) the licensee develop and maintain a

surveillance program to provide reasonable assurance that the containment will be able to 

perform its designed safety function.

The key elements of each RG are somewhat similar with respect to containment geometry but 

with slight differences.

Both RGs require visual inspections but they differ in that RG 1.35 outlines specific areas for the 

visual inspection, in particular the anchor area hardware and the corrosion-preventing filler 

grease. In addition, to the fullest extent possible the ducts containing the tendons should also be

inspected. RG 1.35 also requires that the tendon population be divided into homogeneous 

subgroups consisting of tendons having approximately the same possibility of corrosion and 

similar functions in the overall structural capabilities and properties of the structure.

RG 1.35 requires that containments with ungrouted tendons should be designed so that the 

prestressing anchor hardware is accessible for periodic examination. RG 1.35 also states that 

the inservice inspection should be performed 1, 3, and 5 years after the initial containment 

structural integrity test and every 5 years thereafter.

With respect to the actual testing of the ungrouted tendons, RG 1.35 simply states that the 

tendons selected should be subjected to liftoff or other equivalent tests to monitor loss of 

prestress during each inspection. It is implied that the testing should be of high quality and 

within general industry norms or standards. For example, the testing of the tendons should also 

simultaneously measure the elongation and jacking force with properly calibrated jacks and that 

the maximum test liftoff force should be greater than the maximum design prestressing force. 

RG 1.35 also requires material tests and inspections. Simply stated, previously stressed tendon 

wires or strands from one tendon of each type (i.e., for, “typical” containments, one dome, one 

vertical, and one hoop tendon; for hemispherical dome” containments, one U tendon and one 

hoop tendon) should be removed for testing and examination of corrosion or other deleterious 

effects that may be present. The material testing also includes the checking of the filler grease.

Inservice inspection of prestressed concrete containment structures with grouted tendons, RG 

1.90, is needed to verify at specific intervals that the safety margins provided in the design of 

the containment have not been reduced as a result of operating and environmental conditions. 

The issue with grouted tendons is that there is as yet no real experience to adequately define 

the long-term characteristics of containment structures with grouted tendons. The major concern 

in containment structures with grouted tendons is the possibility that widespread corrosion of the 

tendon steel may occur and remain undetected. The major factors influencing the occurrence of 

corrosion are the: 
3-33

• susceptibility of the tendon steel to corrosion, 

• degree of exposure of the tendon steel to a deleterious environment, 

• extent of temperature variations, and 

• quality of the grout in its installation. 

To address the issues of grout quality the NRC issued RG 1.107 entitled “Qualifications for 

Cement Grouting for Prestressing Tendons in Containment Structures”. The main concern here 

is that the mechanism of corrosion in all conditions is not fully understood. Since many factors 

can influence the development of corrosion or stress corrosion, there is always an area of 

uncertainty with regard to the corrosion of the tendon steel, and it is necessary to monitor the 

structure in a manner that would reveal the existence of widespread corrosion.

Although measuring prestress force through lift-off testing is a straightforward method for 

monitoring ungrouted tendon systems, this approach will not work for grouted tendons since the 

tendons are firmly grouted in place. For this reason, sensors and NDE techniques must be used 

to monitor grouted prestressing systems, but a present lack of industry experience and 

expertise prevents any definitive comment on the use of instrumentation for concrete, i.e., strain 

gauges, stress meters, and strain meters, and their reliability for providing such information. 

When instrumentation that can either be recalibrated or replaced in case of a malfunction or is 

proven to be sufficiently reliable is developed, monitoring the prestress level would be a 

desirable means of assessing the continuing integrity of prestressed concrete structures with 

grouted tendons and would be directly comparable to the lift-off force method for ungrouted 

tendons. In addition, when monitoring the structure at specific locations, it is possible to miss 

damage at a location that is not being monitored, thus missing potential risks to the overall 

health and safety of the structure. As such, other approaches to monitoring should be used in 

addition to monitoring prestress level.

A second means of monitoring the functionality of the containment structure would be to subject 

it to a pressure test and measure its behavior under pressure. RG 1.90 provides two acceptable 

alternative methods of inspecting containment structures with grouted tendons. The first 

method includes monitoring and a longer pressure testing interval and the second method 

includes only pressure testing at a shorter interval. Specifically the two methods are: (1) an 

inservice inspection program based on monitoring the prestress level by means of 

instrumentation plus pressure testing at not exceeding 10 years interval after pressure testing at 

1, 3 and 5 years, and (2) an inservice inspection program based on pressure-testing at time 

intervals not exceeding 5 years after initial pressure testing at 1, 3 and 5 years.

The inservice inspection program recommended by RG 1.90 consists of:

1) Force monitoring of ungrouted test tendons;

2) Monitoring performance of grouted tendons through evaluation:

o prestress levels, or 

o deformations under pressure; and

3) Visual examination.

A brief explanation of the RG 1.90 inservice program is provided to expand on the intent of NRC 

in developing this RG.

The NRC requires the force monitoring of various ungrouted tendons if applicable: three vertical 

tendons, three hoop tendons, three dome tendons, and 4 gamma tendons. These tendons are 
3-34

to be protected with grease instead of grouting, but otherwise are to be identical to the grouted

tendons in the structure. No guidance is provided as to the location of these tendons. This is 

because the monitoring is not meant to capture changes in the structure due to environmental 

or physical effects (with respect to corrosion). The monitoring is instead meant to assess extent 

of concrete creep and shrinkage and relaxation of the tendon steel.

This measurement is necessary to provide a means of verifying the design assumptions in the 

PCCV. Specifically, there is a need to measure volumetric changes in the structure and 

relaxation of the prestressing tendons. Special attention should be paid to this data during a liftoff test, with variation from expected data signifying the need to determine the cause, whether 

corrosion or flawed estimations via calculations. This monitoring can also be used to assess the 

structural response and health following events of interest that have the potential to cause 

damage to the structure. 

The NRC provides two alternatives for monitoring the deformation and prestressing of the 

PCCV structure. Alternative A in RG 1.90 monitors both the prestressing level and the 

deformation of the structure, whereas Alternative B requires measuring the deformation of the 

structure under pressure tests.

In Alternative A, the prestress level is monitored at certain strategically located sections in the 

containment as recommended in RG 1.90. The location of the sensors is meant to 

representatively monitor hoop tendon, vertical tendon, dome tendon, and gamma tendon 

prestressing levels. The licensees are meant to monitor the prestressing levels using a variety 

of sensors, measuring strain, stress, temperature, pressure, and more if available in the 

concrete, rebar, and tendons. 

This sampling is meant to detect degradation in the vicinity of the instrumented section. As 

such, the more sections monitored the better because if corrosion occurs in locations away from 

the instrumented sections, gross degradation would have to occur before the instrumented 

sensors would pick up the degradation. It is especially important to instrument those sections 

with a higher likelihood of corrosion, but it is equally important to remember that corrosion can 

occur anywhere in the structure due to variations in quality of construction and the inability to 

completely control quality construction and variation of in-situ conditions. This method has the 

potential to provide a valuable tool in assuring the safety of the PCCV structures. It can allow 

detection of local degradation prior to serious health problems in the structure that have the 

potential to cause release of radionuclides or the risk of a catastrophic failure of the PCCV. It is 

also a valuable means of monitoring the structure because it allows for continuous monitoring of 

the structure, assessing the health of the structure through-out its life, instead of just at set 

intervals. 

Alternative B requires the licensees to monitor the deformation of the structure under pressure 

loading at set intervals of the PCCVs life. Specific recommendations are made as to the location 

of measurements of deformation to accurately assess the functionality of the structure. This 

method can detect any significant decrease in the stiffness of the structure by measuring the 

elastic response of the structure. Loss of prestress would result in cracking of the structure 

under pressure loads. This method of monitoring allows the licensee to assess the functionality 

of the structure, but can only assess damage once it has become significant enough to affect 

the overall response of the structure. This method does not allow real time monitoring of the 

structure. 

In practice most of the posttensioned containment structures in the nuclear power plants use 

un-grouted tendons. The NRC’s Inspection Manual, Inspection Procedure 62003 entitled 

“Inspection of Steel and Concrete Containment Structures at Nuclear Power Plants” delineates 
3-35

the requirements for the inspection of ungrouted tendons in prestressed concrete containments. 

The NRC has issued information notices which were meant to alert the licenses of potential 

degradations in the respective containment structures. RG 1.35, entitled “Inservice Inspection of 

Ungrouted Tendons in Prestressed Concrete Containments”, and the ASME Boiler and 

Pressure Vessel Code Section XI, Subsection IWL; are useful resources in inspecting for 

degraded post-tensioned tendons in prestressed concrete containments. 

Additionally, it is expected that licensees would use the containment monitoring programs which 

they develop to comply with the requirements of 10 CFR 50.55a to satisfy the monitoring 

requirements of 10 CFR 50.65 (maintenance rule). The NRC encourages licensees to utilize 

such programs for implementing the requirements of 10 CFR 50.65. For those licensees that 

choose to use this program to meet the maintenance rule monitoring requirements, the NRC 

inspector should verify that the performance or condition of the containment structure is properly 

evaluated per the licensees’ maintenance rule program. Guidance for monitoring structures per 

the maintenance rule can be found in RG 1.160 entitled “Monitoring the Effectiveness of 

Maintenance at Nuclear Power Plants.

3.2.2.3 Regulatory Practice in Various Countries

A list of the various regulatory agencies, as well as the regulatory codes related to containment 

and to prestressing tendons where available is listed below in Table 3-3. Table 3-4 identifies the 

reactor types and the prestressing systems used for various countries around the world. Table 

3-5 identifies the inspection methods and acceptance criteria relevant to prestressing systems 

for various countries around the world and Table 3-6 summarizes required inspection frequency 

and relevant code documents for the various countries. The information found in Table 3-3

through Table 3-6 can be found in [Romania 1996, STUK 1996, ACT 1997, CNEN 1997, State 

1997a, State 1997b, State 1998, State 1999, STUK 1999, Stuller, Brandejs et al. 1999, CNEN

2000, STUK 2000, Pakistan 2001, STUK 2001, CNEN 2002, Pakistan 2002, STUK 2002, 

Pakistan 2003, STUK 2003, SG 2004 , SG 2005 , ETC-C 2006, AR 2007, SG 2008, Pakistan 

2008, STUK 2008a, STUK 2008b, STUK 2008c, PP 2010, Pakistan 2010, PP 2011a , PP

2011b, SG 2012 , Licensing 2012 , Pakistan 2012, Regulatory Guide 1.190 2012].
3-36

Table 3-3 Regulatory agencies and codes/guidance.

Country Regulatory 

agency

Regulatory 

code related to 

containment 

vessel and 

testing

Standards for 

tendons (grouted)

Standards for 

tendons (ungrouted)

Argentina ARN AR 3.4.3, AR 3.8.1

3.2.3 No specific standard. AR 3.4.3 requires that 

containments be designed to withstand accident 

loads and that the systems need to be monitored. 

AR 3.8.1 preliminary testing and commissioning of 

nuclear reactors power. Commissioning of a 

Nuclear Power Plant must be considered initiated 

with the load of the fuel and moderator at the 

Reactor Pressure Vessel; the preliminary test, 

however, fall within the construction phase of the 

installation.

The preliminary testing program must demonstrate 

that such tests will verify compliance on the part of 

the components, equipment and systems 

installation that apply at this stage, of the 

requirements and design goals incorporated in the 

safety report.

The Start-up program must allow demonstrate that 

the design objectives will be met in all 

components, equipment and systems that the 

facility will operate safely in both normal operation 

and in operational occurrences and that the 

systems intended to deal with accidental situations 

will function properly.

Armenia

Armenian 

Nuclear 

Regulatory 

Authority

Uses Russian 

Rules and Norms See Russia See Russia

Belgium

Federal Agency 

for Nuclear 

Control

GRR-2001

Nuclear Power is being phased out. No new plants 

can be built, and all plants are to be 

decommissioned after 40 years of service

Bulgaria

Nuclear 

Regulatory 

Agency of the 

Republic of 

Bulgaria

Act on the Safe 

Use of Nuclear 

Energy, Nuclear 

Safety, and 

Radiation 

Protection

It is up to the licensee to provide plans for testing 

and design and demonstrate safety and 

adequacy. The regulator must approve the 

licensee’s plans.

Canada

Canadian 

Nuclear Safety 

Commission

RD-337 (design)

RD-334 (aging 

management)

CSA N287 Series

China

National Nuclear 

Safety 

Administration 

China Atomic 

Energy Authority

HAF 0200 Based on USNRC regulatory Environment as of 

1991, no updates to date. (Use US RG 1.90)

Czech 

Republic

State Office for 

Nuclear Safety

Decree of the 

SUJB No. 

195/1999

309/2005

No grouted tendons used, only ungrouted 

tendons.

No national standards. Design is Russian, later 

modification according US practice have been 

done.
3-37

Table 3-3 Regulatory agencies and codes/guidance (continued).

Country Regulatory 

agency

Regulatory code 

related to 

containment vessel 

and testing

Standards for 

tendons 

(grouted)

Standards for 

tendons 

(ungrouted)

France ASN (Nuclear 

Safety Authority) RCCG 88 Tome 2 ETC-C for AREVA 

EPR Design

Germany

Federal Ministry 

for the 

Environment, 

Nature 

Conservation 

and Nuclear 

Safety

Safety Standards of the 

Nuclear Safety 

Standards Commission

Plant Specific Plant Specific

Hungary

Hungarian 

Atomic Energy 

Authority

Nuclear Safety 

Regulations per Gov. 

Decrees 118/2011 

(VII.11.) and 37/2012. 

(III. 9.)

No prestressing

India

Atomic Energy 

Regulatory 

Board (AERB)

French code RCC-G for 

concrete containment 

without metallic liner

ASME Section-III Divisio2 for containment with 

metallic liner

AERB/NPPPHWR/SG/D-21, 2008 

for Containment System 

Design

AERB/NPP/SG/O-15, 

2004 for Proof and 

Leakage Rate Testing of 

Reactor Containments

AERB/NPP/SM/CSE-2, 

2004 for Inservice

Inspection of Civil 

Engineering Structures 

AERB/SM/CSE-1, 2002 

for Maintenance of Civil 

Engineering Structures

Plant Specific, 

depending on the 

period of design 

and construction.

Current/new 

designs require 

complete scheme of 

long-term structural 

monitoring.

Plant Specific.

Current/new 

designs require 

complete scheme 

of long-term 

structural 

monitoring.

Japan

Nuclear 

Regulation 

Authority

Law for the Regulation of 

Nuclear Source Material, 

Nuclear Fuel Material, 

and Reactors

Guideline for 

assurance of 

structural integrity 

of concrete vessel 

and inservice

inspection during 

plant life JEAG4203

Grouted Tendons 

Not Used
3-38

Table 3-3 Regulatory agencies and codes/guidance (continued).

Country Regulatory 

agency

Regulatory code 

related to 

containment vessel 

and testing

Standards for 

tendons 

(grouted)

Standards for 

tendons 

(ungrouted)

Korea

Nuclear Safety 

and Security 

Commission

Korea Institute 

of Nuclear 

Safety

Regulation of the Atomic 

Energy Act

KINS/RG-N04.10

Follows US, 

French, and 

Canadian Code

KINS/RG-N04.07 

and USNRC RG. 

1.90

Mexico

National 

Commission on 

Nuclear Safety 

and Safeguards

CNSN-003 No prestressing

Netherlands

Ministry of 

Economic 

Affairs, 

Agriculture and 

Innovation

Department of 

Nuclear Safety, 

Security, and 

Safeguards

Nuclear Energy Act No prestressing

Pakistan

Pakistan 

Nuclear 

Regulatory 

Authority

Pakistan Nuclear Safety 

and Radiation Protection 

Regulations

PAK 911

Proven alternative 

and/or indirect 

methods such as 

surveillance of 

reference items or 

use of verified and 

validated 

calculation methods 

can be used

Follows ASME 

Boiler and 

Pressure Vessel 

Code for Inservice

Inspection Methods

Spain Nuclear Safety 

Council Nuclear Energy Act No grouted tendons Reg Guide 1.35

Sweden

SSM (Swedish 

Radiation Safety 

Authority)

SKI Report 02:58 Information 

unavailable

Follows US Reg. 

Guide 1.35

United 

Kingdom

Office for

Nuclear 

Regulation (An 

Agency of HSE)

The Licensing of Nuclear 

Installations

No Grouted 

Tendons At 

present, but a new 

plant is being 

reviewed

NS-TAST-GD-020 

- Revision 2 which 

references various 

others (SAP, IAEA 

NS-G-1.10, BS 

4975, ACI 318, ACI 

359)

United 

States of 

America

Nuclear 

Regulatory 

Commission

10 CFR 50.55a RG 1.90 & ASME 

section XI, IWL

RG 1.35 & ASME 

section XI, IWL
3-39

A list of the types of reactors, containments, and prestressing systems used in different 

countries is listed below in Table 3-4. 

Table 3-4 Reactors, containments, and pre-stressing systems used around the world.

Country Type of 

reactors Type of containment

Types of 

prestressed 

concrete 

tensioning 

systems used 

in pressure 

vessels

Types 

allowed 

under 

regulatory 

environment

Argentina PHWR Pressure Suppression, Steel None Both

Belgium PWR

Steel/Concrete, Concrete/Liner, 

Double Concrete, Prestressed 

Concrete

Grouted Not Specified

Canada PHWR Reinforced Concrete, Prestressed 

Concrete

Grouted and 

Ungrouted Both

China FBR, PWR,

PHWR Prestressed Grouted Grouted

Czech 

Republic PWR Reinforced Concrete, Prestressed 

Concrete Ungrouted Both

Finland PWR, BWR Ice Condensed, Concrete, 

Prestressed Concrete

Grouted and 

Ungrouted Both

France PWR Prestressed Concrete, Reinforced 

Concrete, Steel

Grouted and 

Ungrouted Both

Germany PWR, BWR

Steel, Reinforced Concrete, 

Pressure Suppression, Dry, 

Prestressed Concrete, Chemical 

Shim

Grouted and 

Ungrouted Both

Hungary PWR

Reinforced Concrete Steel Lined 

Containment With Pressure 

Suppression System (Bubbler 

Condenser Tower)

None

India PHWR, PWR, 

FBR, BWR

Prestressed Concrete, Reinforced 

Concrete, Pressure Suppression

Grouted and 

Ungrouted Both

Japan BWR, PWR Steel, Prestressed Concrete, 

Reinforced Concrete Ungrouted Ungrouted

Korea PWR, PHWR Steel With Reinforced Concrete

And Prestressed Concrete

Grouted and 

Ungrouted Both

Mexico BWR GE Mark II, Pressure Suppression None

Netherlands PWR Pressure Suppression None

Pakistan PWR, PHWR Steel/Concrete, Reinforced 

Concrete Ungrouted Not Specified

Russia PWR, FBR, 

LWGR

Reinforced Concrete, No Shell, 

Steel, Prestressed Concrete Ungrouted Not Specified

Spain PWR, BWR

Reinforced Concrete, Prestressed 

Concrete, Pressure Suppression, 

Steel

Ungrouted Not Specified
3-40

Table 3-4 Reactors, containments, and pre-stressing systems used around the world 

(continued).

Country Type of 

reactors Type of containment

Types of 

prestressed 

concrete 

tensioning 

systems used 

in pressure 

vessels

Types 

allowed 

under 

regulatory 

environment

Sweden BWR, PWR

Prestressed Concrete, Pressure 

Suppression, Steel, Reinforced 

Concrete

Grouted and 

Ungrouted Both

Ukraine PWR Prestressed Concrete, No Shell, 

Reinforced Concrete Ungrouted Not Specified

United 

Kingdom GCR, PWR

Prestressed Concrete, Steel, 

Pressure Suppression, Concrete, 

Bio-Shield, Vented

Ungrouted

Ungrouted, 

but in the 

process of 

reviewing 

grouted

United 

States of 

America

PWR, BWR Steel, Prestressed Concrete, 

Reinforced Concrete

Grouted and 

Ungrouted Both

A comparison of the different regulatory practices related to grouted and ungrouted tendons is 

shown below in Table 3-5 and Table 3-6.
3-41

Table 3-5 Inspection methods and acceptance criteria for various countries.

Country Type of 

tendon Inspection method Acceptance criteria

Canada

Grouted

. Instrumented monitoring

OR

. Test beams (a) Lift-Off tests on at least 

four ungrouted test beam that matches 

material properties of the containment 

structure. (b)Destructive Test on one test 

beam to examine the condition of the 

tendon and corrosion protection medium. 

(c)Flexural tests carried out on grouted

test beams to check concrete cracking.

AND

. Visual Inspection

Measurements confirm that structure 

behaves as designed and there is no 

unusual trend 

OR

(a) Measured prestressing force is 

considered acceptable if it is higher 

than the value predicted in design, 

taking into account long term 

prestressing losses; (b) Any sign of 

tendon corrosion is considered 

unacceptable; and (c) Beams are 

required to be loaded to 95% of the 

design cracking moment. Results 

are acceptable if no cracking of the 

concrete takes place under the test 

load.

AND

No degradation 

Ungrouted

Lift-Off tests (to assess prestress force) 

on randomly selected samples from each 

tendon group in the structure - 4% of 

tendon - population with minimum 4 and 

maximum 10 per type (if previous tests 

successful can be 2% of population, with 

a minimum of 3 per type and maximum 

of 5 per type

. Inspection of the tendon wires or strands 

with visual inspection for corrosion and 

defects, ultimate tensile strength and 

elongation - a minimum of one tendon for 

each group shall be de-tensioned and a 

wire or strand shall be removed for visual 

inspection and laboratory testing 

. Inspection of the tendon corrosion 

prevention medium

. Inspection of tendon anchor regions

Measured tendon force is 

considered acceptable if it is larger 

than the predicted value taking into 

account time dependent losses due 

to creep, shrinkage, and relaxation 

Tendon wire or strand is considered 

acceptable if the test results are not 

less than the minimum specified 

values 

Acceptable if laboratory test results 

are found to be within the original 

specifications described is CSA 

N287.2.

Acceptable if no abnormalities are 

detected. 
3-42

Table 3-5 Inspection methods and acceptance criteria for various countries 

(continued).

Country Type of 

tendon Inspection method Acceptance criteria

China Grouted

Force monitoring of ungrouted test 

tendons (3 vertical, 3 hoop, 3 dome)

Monitoring performance of grouted 

tendons by (a) Monitoring of prestress 

level or( b) Monitoring of deformation 

under pressure

Visual examination

Verify forces against design 

assumptions, must be above 

acceptable band 

(a) Average prestress force must be 

within acceptance band and 

prestress force must be above the 

design prestress force; 

(b) Deformation measured under 

the maximum test pressure at any 

location cannot increase by more 

than 5% of that measured during 

the ISIT under the same pressure 

No crack patterns observed at 

structurally critical areas indicating a 

significant decrease in the spacing 

or an increase in the widths of 

cracks compared to the zero 

pressure, no degradation of the 

anchor hardware, no degradation of 

protection of anchor hardware

Czech

Republic Ungrouted

Concrete: 

. Visual inspection of overall vessel with 

respect to cracks, degradation and 

defect, corrosion of rebar

. Detailed visual inspection for selected 

areas with respect to crack width and 

length, non-destructive tests of concrete 

strength

Liner: 

. Visual inspection of surface 

. Non-destructive measurement of 

thickness

Tendon: 

. Visual inspection of anchorage area 

(presence of water, corrosion, damage)

. Lift-off tests 

. Removal of tendon for detail visual 

inspection and testing tensile strength, 

yielding strength, and elongation

Measurement: 

. Temperature corrected vibrating wire 

strain gauge - deformation and 

temperature of concrete

. Vibrating wire extensometers – prestressing force in anchors

Concrete: if newly developed 

cracks, degradation, and defects 

are found, the structural integrity of 

vessel must be assured by 

investigating the cause and taking 

proper remedial actions if required

Liner: no cracks, corrosion or 

damage, thickness meets 

requirements

Tendon: prestressing force 

consistent with design assumptions. 

Tendons and anchors free from 

corrosion, no physical defect, and 

mechanical properties satisfy the 

specified values. 

Measurement: deformation and 

prestressing force consistent with 

design assumptions

Finland Grouted

. Four vertical and hoop tendons left 

ungrouted for monitoring 

. Fiber optic sensors applied to grouted 

tendons to monitor strain

Tendons are only monitored for 

problems. 
3-43

Table 3-5 Inspection methods and acceptance criteria for various countries 

(continued).

Country Type of 

tendon Inspection method Acceptance criteria

France Grouted

Based on:

. Measurement by instrumentation of 

strains in concrete and evaluation of the 

resulting losses in prestressing force

. Periodical pressure tests as DBA level 

(Pa) including complete instrumentation 

measurements and visual inspection

. Monitoring of some tendons left 

ungrouted (on prototypes only)

Permanent observation of 

instrumentation measurements to 

check that with time:

. No abnormal strain is observed

. Resulting prestress forces are within 

predicted values

. Observation of elastic behavior 

during pressure tests and limited 

change in deformations

Germany Grouted

. Leak rate measured periodically at half 

the design pressure

. Visual inspections of the containment on 

a yearly basis

< 0.25% per day leak rate for PWR 

and 

< 1.0% per day leak rate for BWR

India

Grouted

and 

Ungrouted

. Monitoring of the state of pre-stress 

indirectly by measuring the strain in 

containment wall/dome using embedded 

strain gauges. 

. Keeping a few number of ungrouted

cables with load cells fitted in the cable 

anchorage assembly of real time 

monitoring of pre-stress level

. Keeping a few number of ungrouted

cables with provision for periodic 

measurement of prestress by lift-off 

technique 

. Periodic measurement of resistance 

value of cables to calculate corrosion 

rate

Evaluation of collected data and 

assessment of structural safety visà-vis design requirements.

Japan Ungrouted

Concrete:

. Visual inspection of overall vessel with 

respect to cracks, degradation, and 

defect 

. Detailed visual inspection for selected 

areas with respect to crack length, 

extent and circumstance of degradation 

and defect

. Corrosion of rebar

Tendon: 

. Visual inspection of tendon and 

anchorage lift off or equivalent

. Removal of sample tendon for testing 

tensile strength, yielding strength, and 

elongation

Concrete: if newly developed 

cracks, degradation, and defects 

are found, the structural integrity of 

vessel must be assured by 

investigating the cause and taking 

proper remedial actions if required

Tendon: prestressing force beyond 

the effective design force. Tendon 

free from corrosion, no physical 

defect, and mechanical properties 

satisfy the specified values. 

Anchorage free from corrosion, no 

local damage, and no unusual 

cracks observed in its peripheral 

concrete
3-44

Table 3-5 Inspection methods and acceptance criteria for various countries 

(continued).

Country Type of 

tendon Inspection method Acceptance criteria

Korea Grouted

At present, two different systems are used to assess plants in Korea. The 

Wolsong units follow: the Canadian regulatory practice defined in CSA-N287.7, 

US regulatory guidelines set forth in US RG 1.90, and FSAR. At the Ulchin 

plant, inspection procedure follows RCC-G, ASME Sec XI (visual inspection), 

and FSAR. Please see US and Canadian sections for description of testing 

requirements.

Korea Ungrouted

. Visual examination

. Tensile force monitoring test (Lift off or 

equivalent tests performed on randomly 

selected tendons)

. Test and inspection of tendon material 

(Removal of one tendon of each type)

. Inspection of the grease filler

. Prestress forces within the predicted 

limits.

. Tensile strength of the sampled 

tendon shall satisfy the minimum 

requirements of the material.

Spain Grouted See US requirements See US requirements

United

States

Ungrouted

. Based on inspection of a selection of 

tendons of each type (vertical, hoop, 

dome) (about 4% of total number)

. Visual inspection of anchorage system

. Lift off or equivalent tests performed on 

selected tendons

. Removal of one tendon of each type for 

material testing of steel and grease.

. No pressure test associated with the 

structural surveillance tests.

. Prestress forces within the predicted 

limits

. Not more than one tendon defective 

in total sample population as long 

as adjacent tendons acceptable

. Tensile failure tests on removed 

tendon higher than UTS

Grouted 

(Alternative 

A)

. Monitoring by instruments and sensors 

to predict losses in tendon force

. Completed by visual inspection and lift 

off tests on tendons left ungrouted (3 

vertical, 3 hoop, and 3 dome) 

. No pressure test associated with testing

Verification that the evaluated 

prestress forces are within the 

predicted region, as well as the 

forces of tendons left ungrouted.

Grouted 

(Alternative 

B)

. Pressure tests at least at DBA level (Pa) 

and comparison of the deformation of 

the containment with ISIT 

measurements

During pressure test: 

. Limited increase of deformations 

compares with ISIT (less than 5%)

[REGULATORY GUIDE 1.90 2012]

. Limited increase of cracks

. Verification that the forces in 

tendons left ungrouted are 

acceptable.
3-45

Table 3-6 Frequency of inspection and codes related to testing for various countries.

Country Type of 

tendon Frequency of tests Code

Canada

Grouted 6 years CSA-N287.7

Ungrouted

6 years, but if level of degradation is 

insignificant can be increased up to 12 

years

CSA-N287.7

China Grouted

1, 3, and 5 years after initial structural 

integrity test, and every 5 years 

thereafter, alternative (b) follows pressure 

testing

NRC Reg Guide 1.90

Czech 

Republic Ungrouted

Visual inspections every year for first 5 

years. After that, inspection of concrete 

and liner every year and inspection of 

tendons every sixth year. Measurement 

every month.

No national codes. Tests are 

specified in design, based on mix of 

Russian and US practice.

France Grouted

. Frequent instrumentation measurements 

(continuous since installation for the 

latest containments)

. Pressure tests at: first shut-down for 

refueling, every 10 years

RCCG 88 Tome 2

India

Grouted 

and 

Ungrouted

As specified in safety guides and 

manuals

. AERB/NPP/SG/O-15, 2004 for Proof 

and Leakage Rate Testing of 

Reactor Containments

. AERB/NPP/SM/CSE-2, 2004 for 

Inservice Inspection of Civil 

Engineering Structures

Japan Ungrouted

1 yr, 3 yr, and then every 5 yr after ISIT. 

The frequency can be reduced if the if 

the same type of plant exists on the same 

site.

Guideline for assurance of structural 

integrity of concrete vessel and 

inservice inspection during plant life 

(tentative) JEAG4203

Korea

Grouted

Two different systems are used to assess 

plants in Korea. The Wolsong units 

follow: the Canadian regulatory practice 

defined in CSA-N287.7, US regulatory 

guidelines set forth in US RG 1.90, and 

FSAR. At the Ulchin units 1&2, inspection 

procedure follows RCC-G, ASME Sec XI, 

and FSAR. Please see US and Canadian 

sections for description of testing 

requirements.

CSA-N287.7

US RG1.90

ASME Sec XI, Div. 1

FSAR

Ungrouted At 1, 3, and 5 years, and then every five 

years after ISIT

KINS/RG-N04.07

USNRC RG. 1.90
3-46

Table 3-6 Frequency of inspection and codes related to testing for various countries 

(Continued).

Country Type of 

tendon Frequency of tests Code

Spain Ungrouted See US Ungrouted NRC Reg Guide 1.35,

ASME XI IWL

Sweden

Grouted No special requirements for grouted 

tendons SKI Report 02:58

Ungrouted See US Ungrouted SKI Report 02:58 (follows US Reg 

Guide 1.35)

United 

States

Ungrouted At 1, 3, and 5 years, and then every 

five years after ISIT

10 CFR 50

NRC Reg Guide 1.35

Grouted 

(Alternative 

A)

Instrumentation readings every two 

months from installation to first 

inspections to ISIT, then every 6 

months. Readings under pressure at 1, 

3, 5, and then every 10 years after 

ISIT.

10 CFR 50

NRC Reg Guide 1.90

Grouted 

(Alternative 

B)

Readings under pressure at 1, 3, 5, 

and then every 7 years after ISIT.

The following subsections provide additional detail describing regulatory practice in various 

countries.

3.2.2.4 Canada

Canada also has experience with the use of grouted tendons in their CANDU reactors 

(specifically the CANDU 6). Within the Canadian nuclear codes (CSA 287 series), there are 

multiple clauses related to the testing and inservice examination required for grouted tendons. 

CSA N287.6-11 requires that instrumentation be provided to evaluate the behavior of the 

structure and the actual stress values during pre-operational proof testing. In addition, 

instrumentation is to be provided to be used for the life of the plant, to monitor deformation of 

the containment structure during leakage rate testing (to ensure elastic behavior of the 

containment structure), and to verify the integrity of the prestressing system. Furthermore, CSA 

N287.7-08 requires that if prestressing is used as the principal reinforcement in the concrete 

containment structures, then it shall be subject to an integrity evaluation, assessing time related 

factors like shrinkage and creep of concrete, stress relaxation, and deterioration. For these 

evaluations, they recommend vibrating wire strain gauges, fiber optic sensors, and 

thermocouples. In addition, instrumented test beams are used to test the tendons, with flexural 

tests, lift-off tests and inspections performed under the same conditions as the containment.

For new design, Canadian regulators are considering the instrumentation and monitoring of the 

structures to ensure detection of cracks, displacements, strains, and vibrations. 

3.2.2.5 China

At present, China has 16 nuclear power plants, all but two of which are PWRs (other two are 

PHWRs). China also has 26 plants currently under construction. As of 2002, the Chinese 
3-47

nuclear regulatory environment is defined in HAF 0200(91), “Code on the Nuclear Power Plant 

Design” (China NNSA, 1991), which is based on the USNRC regulatory environment at that 

time. All PCCV nuclear plants in China had grouted tendons in 2002, and were to be monitored 

using US RG 1.90 [Regulatory Guide 1.90 2012]. To date, this regulatory environment has not 

been changed (as per International Atomic Energy Agency (IAEA) Power Reactor Information 

System (PRIS)). 

3.2.2.6 Czech Republic

At present, only ungrouted tendon systems are used in the Czech Republic. 

3.2.2.7 Finland

Finland has multiple prestressed concrete containment structures with grouted tendons. Both 

Olkiluoto Unit 1 and Olkiluoto Unit 3 (OL3) have grouted tendon systems. For monitoring of the 

PCCVs with grouted tendons, periodic leakage tests are performed in which strain and 

temperatures are recorded at intervals of 0.05 MPA. The containments are equipped with strain 

and temperature sensors, and leakage is observed and assessed versus ANSI N45.4 (for leak 

rate testing, for US plants, 10 CFR Part 50, Appendix J for leak rate testing and is mandated by 

rule) [ETC-C 2006]. During these tests, the possibility of corrosion is evaluated from available 

test data, and visual crack inspections are performed. OL3 is a new construction, and the 

tendons are being equipped with fiber optic sensors for monitoring of strain in grouted tendons. 

As mentioned previously, partial and full mock-ups of grouting practice are being tested to 

ensure high quality grouted tendons are produced. In addition, four horizontal tendons and four 

vertical tendons are being left ungrouted for stress monitoring.

3.2.2.8 France 

France has provisions within its regulatory environment for the use of grouted tendons in 

nuclear containment vessels. For the construction of the PCCV, European standards are used, 

and the French certification body for prestressing (ASQEP) is used. Grout and injection 

methods are controlled by the European Standards (EN 445 to EN 447). Of note is the fact that 

grouting is performed in two phases for those vertical tendons which go through the dome, and 

single phase grouting is used for pure vertical or horizontal tendons. In addition, full scale mockup testing is required to be performed to ensure the grout injection procedure is adequate.

The requirements of France are similar to those of Alternative A for grouted tendons in the 

United States. France requires the instrumentation and measurement of strains in concrete and 

the evaluation of the resulting losses in prestressing force. To assess the shrinkage and creep 

of the concrete, France requires the measurement of global lengthening, diameter variation, and 

local deformation using vertical invar wires, pendulums, and extensometers throughout the 

lifespan of the PCCV. France requires periodic pressure tests of the containment at DBA level 

(Pa). In addition to the periodic testing, France requires complete instrumentation 

measurements and visual inspection. France requires four vertical tendons to be left ungrouted 

and fitted with dynamometers to monitor transfer losses and tension losses throughout the 

design life of the nuclear power plant. 

3.2.2.9 Japan

Japan at present does not allow for grouted tendons in their plants, so all requirements are 

related to the monitoring of ungrouted tendons. 
3-48

3.2.2.10 Korea

At present, the Republic of Korea has 21 operating nuclear power plants, comprised of 17 

PWRs, and 4 CANDU pressurized heavy water reactors. In addition, there are 7 units under 

construction. The plants Ulchin Unit 1 and 2, and Wolsong Units 1, 2, 3, and 4, have grouted 

tendon systems. Each of the aforementioned plants has slightly different inspection 

requirements. As part of the inservice inspection routine required by the Korean Institute for 

Nuclear Safety (KINS) for the Wolsong Units, the concrete surface is inspected, indirect 

prestress force tests using the test beam method, integrated leak rate tests, and local leak rate 

tests are performed. The inspection procedure is to follow Can/CSA-N287.7 (CSA, 1980), US 

RG1.90, and FSAR (KHNP, 1978-2005). At Ulchin, the liner plate is inspected, as well as leak 

resistance and strength tests are performed. For Ulchin, the inspection procedure is to follow 

RCC-G Part 3 (AFCEN, 1981), ASME Sec XI, Div. 1 IWE, and FSAR [Park and Hong 2009, 

Park 2010]. The Ulchin units were connected to the grid in 1988 and 1989 respectively, and the 

Wolsong units 1-4 were connected to the grid in 1982, 1997, 1998, and 1999 respectively. No 

reported corrosion problems have been observed in these tendon systems

At present, the Korea Electric Power Industry Code (KEPIC) is under development by the Korea 

Electric Association (KEA) to be used as the comprehensive code detailing inspection 

requirements. The last edition was released in 2005, and the KEPIC website states that new 

editions come out every five years. This document is in Korean, and the authors have been 

unable to find a translation at present, or a more recent version. This addition has yet to be 

implemented by KINS.

3.2.2.11 Sweden

At present, both ungrouted and grouted tendon systems are used in nuclear power plants in 

Sweden. Sweden has four containments with grouted tendons, Ringhals 1, and Oskarshamn 1 

through 3. In addition, Sweden has six containments with ungrouted tendons, Ringhals 2, 3, 4 

and Forsmark 1 through 3. Ungrouted tendons are monitored in accordance with US Reg Guide 

1.35. There are no special regulations for testing tendons for Swedish containments with 

grouted tendons [Anderson 2005, Anderson, Hansson et al. 2008]

3.2.2.12 United Kingdom

In the United Kingdom, grouted tendons are not used in containment vessels. There is concern 

within the UK nuclear industry about relaxation due to temperature and about being unable to 

detect tendon failures when grouted tendons are used. However, there has been an increase in 

non-nuclear grouted tendon projects in the UK. Also a design for the UKEPR containment 

design using grouted tendons is currently under review. This suggests that the UK may soon 

address the issue of grouted tendons for nuclear power plant applications.

3.2.2.13 Summary of Comparison of French, Canadian, and United States Regulatory 

Practice Concerning Grouted Tendons

Both the United States and Canada use the ASME Boiler and Pressure Vessel Code (BPVC) as 

the starting point for the design, inspection, and monitoring of the pressure vessels and 

prestressing systems. The Canadian standards provide additional requirements where needed 

because of the unique nature of the plants used in Canada. The N287 series of requirements 

covers the requirements for concrete containment systems in Canada. In the United States, 

additional guidelines are defined in the US RGs 1.35 and 1.90 for ungrouted and grouted 

tendon systems respectively. The US and Canadian standards are used frequently by other 

countries as either a starting point for regulation and additional rule making or directly as the 
3-49

standard by which plants are monitored. The French regulatory guide specifically for grouted 

tendons at EPR plants is the EPR Technical Code for Civil Works (ETC-C), which is being used 

for comparison in this section. It is important to note that the ETC-C was developed specifically 

for the AREVA/EDF EPR design, and is not directly applicable to other reactor or containment 

designs. 

Some of the major similarities between the US, Canadian, and French codes are:

• Alternative A of the US system directs and the French system requires sample tendons 

to be left ungrouted in the concrete containment so that prestressing forces can be 

evaluated. The Canadian code requires that prestressing force be evaluated in test 

beams that have been left ungrouted.

• Alternative A of the US system directs and France requires instrumented monitoring to 

evaluate prestressing force.

• Alternative A and B of the US system directs and France requires pressure tests on the

structure with subsequent monitoring of the structural response.

Some of the major differences between the three countries’ codes are:

• The ETC-C requires continuous monitoring of creep and shrinkage during construction 

as well as during operation. There are no such requirements for continuous monitoring 

of creep and shrinkage effects in ASME BPVC, however US RG 1.35.1 provides 

guidance for monitoring of creep and shrinkage losses at prescribed intervals.

• ASME BPVC provides specific requirements for monitoring cracks during testing while 

the ETC-C does not.

• The ETC-C provides provisions for conducting both local and overall (integrated) leak 

tests and a resistance (structural integrity) test, at the same time. ASME BPVC specifies 

the requirements for a Structural Integrity Test while the overall leak test or integrated 

leak rate test requirements for U.S. containment vessels are specified by the US NRC in 

10CFR50, Appendix J, which references ANSI test procedures ANSI/ANS N45-2-1974 

and ANSI/ANS N56.9-1987.

• The Canadian codes require the testing of individual sample test beams with grouted 

and ungrouted tendons.

• Canada is in the process of updating current testing procedures, and is investigating the 

use of incorporating fiber optic sensors into the grouted tendons for continuous strain 

monitoring.

3.2.3 PCCV Inservice Inspection Methods Related to Post-Tensioning

A number of inspection methods are discussed below which have the potential to assist in 

assessing the prestressing state of grouted tendons in a PCCV. NDE methods have the 

potential to provide valuable insight to the overall health of the PCCV structure, and their 

development and use should be considered. Additionally, it is important to note that there is no 

one evaluation technique that is appropriate for all conditions, rather the methods described in 

the subsections below should be viewed as a set of tools that should be used in conjunction 

with one another to develop a more thorough assessment. 
3-50

3.2.3.1 Visual Inspection

A key area of monitoring the health of PCCV structures recommended by the NRC in 

conjunction with the monitoring mentioned previously is visual inspection. Visual inspection has 

proven to be an invaluable method for detecting structural damage in the past, and continues to 

provide valuable insight to investigators assessing structural health. Visual inspection is 

especially recommended in those areas with structural discontinuities or heavy stress 

concentrations. Examination of the concrete and tendon anchorage in particular is very 

important and useful in assessing structural health. 

Every country which uses grouted tendons in PCCVs requires visual inspection (see Section 

3.2.2.3). The timing of these inspections varies slightly by country, depending on leak rate test 

schedules and the periodicity of visual inspection that each country has specified. When 

performing visual inspection, investigators are looking for evidence of corrosion, cracking of 

concrete, damage to the anchoring system, or damage to concrete or exposed rebar [ASME

2010b]. 

3.2.3.2 Acoustic Monitoring

Acoustic monitoring uses locally placed sensors around a structure to detect acoustic emissions 

that are produced when a material fails. This monitoring can detect material damage that is 

invisible to the naked eye (such as microscopic flaws or internal damage beneath the visible 

surface). This method requires continuous monitoring of the structure. It cannot detect damage 

that has already occurred, but rather can be used to detect changes in the current structure. 

Standard practice for characterizing the instrumentation needed for acoustic monitoring is 

defined in ASTM E750 [ASTM 2010].

Acoustic monitoring has successfully been used to detect tendon wire breaks, rebar breaks, 

concrete cracking and crushing, and liner tearing and leakage. In the 1:4 Scale prestressed 

concrete containment test mentioned in Section 2, acoustic monitoring was used successfully to 

detect all of these failures in a PCCV with ungrouted tendons [NUREG/CR-6810 2003]. An 

example of the monitoring of tendon breaks is shown in Figure 3-9. This plot shows the energy 

levels calculated using the acoustic data, the spikes in energy correspond to the breaking of 

tendons, and the corresponding energy which is released in that tendon break.
3-51

Figure 3-9 Energy vs. time of wire breaks in 1:4 scale SMFT [NUREG/CR-6810 2003]

In addition to success with ungrouted tendons, acoustic monitoring has successfully been used 

to detect damage in grouted tendon systems. Work done in France has shown the potential of 

using acoustic monitoring to detect stress corrosion cracking in grouted prestressing strands in 

bridges [Perrin, Gaillet et al. 2008, Ramadan, Gaillet et al. 2008]. Acoustic monitoring has also 

been used on a bridge in Switzerland to successfully detect wire breaks in grouted tendons 

[Fricker and Vogel 2007]. The Canadians also have extensive background with the use of 

acoustic monitoring of civil structures and commercial structures [Tozser and Elliott 2000, 

Tozser, Barker et al. 2010]. 

Acoustic monitoring can provide a valuable tool to detect damage in a PCCV, and especially in 

grouted tendons which cannot be monitored using typical methods employed for ungrouted 

tendons. The drawbacks of this method are the need for continuous monitoring, the need for 

qualified operators to monitor and interpret the data being presented, and the inability to directly 

assess the scope of the damage. 

3.2.3.3 Ultrasonic Pulse Echo

The ultrasonic pulse echo method is a non-destructive testing (NDT) method which can be used 

to assess the structural integrity of concrete. This method measures the time of travel of an 

ultrasonic pulse through an object. In general, the higher the velocity of the pulse is through an 

object, the better the quality of the object. For example, when cracks or voids are present in 

concrete, the ultrasonic pulse is unable to travel directly through the concrete in a straight line, 

and the pulses that return to the transducer have slower return times due to the indirect path 

they were forced to take. As such, this method can be used to detect cracks, voids, intrusion, 
3-52

and honeycomb conditions in concrete within approximately 15 inches of the surface of the 

concrete. An example of an ultrasonic system is shown below in Figure 3-10.

Figure 3-10 Ultrasonic inspection setup [Lafhaj, Goueygou et al. 2006]

The overall quality of concrete has been shown to be assessable via the sonic pulse technique 

through operational experience and research. If the pulse velocity is 4,500 m/s (14763 ft/s) or 

greater, the concrete quality is considered good to excellent. If the pulse velocity is below 4,000 

m/s (13123 ft/s) for normal weight concrete, the concrete is of questionable quality [Regulatory 

Guide 1.90 2012]. ASTM C597 [ASTM 2009] details the standard test method for pulse velocity 

through concrete. This method can be combined with the rebound hammer test [ASTM 2013] or 

other NDT methods to locate corrosion prone areas.

In addition to assessing the overall quality of concrete, much research has been done to assess 

the strength, porosity, permeability, water/cement ratio, and detection of voids and cracks in 

concrete as a function of pulse velocity [Phoon, Wee et al. 1999, Rapoport, Popovics et al. 

2000, Nogueira and Willam 2001, Liang and Wu 2002, Matusinović, Kurajica et al. 2004, 

Ramamoorthy, Kane et al. 2004, Lafhaj, Goueygou et al. 2006, Lin, Kuo et al. 2007, Lorenzi, 

Tisbierek et al. 2007, Santhanam 2010]. Some of the general trends observed experimentally 

are:

• as porosity increases, pulse velocity decreases

• as compressive strength increases, pulse velocity increases

• as the w/c ratio increases, pulse velocity decreases

• in general, as moisture content increases, the velocity increases (small influence)

• for high strength concrete (greater than 68.95 Mpa (10,000 psi) fc), additional NDE 

evaluation is recommended to assess the quality of concrete [Santhanam 2010]

Although Ultrasonic inspection shows potential for monitoring material defects close to the 

surface of PCCV structures, it is currently not possible to use this method to assess the 

prestressing level of grouted tendons because of the depth limitations of the ultrasonic waves 

(effective at less than 1.5m (4.9 ft.)). The grouting causes too much energy attenuation for the 

waves to reach the tendons and accurately relay structural health details. Another limitation to 

detecting damage in concrete is the presence of water within a specimen. If cracks are filled 

Transducer

Sample

Air

Buffer SI(t) S2(t)
3-53

with water, it will not be possible for the ultrasonic method to detect the crack in the concrete. In 

addition, the path length, lateral dimensions of the concrete specimen, reinforcing steel, and 

moisture content of concrete can influence the pulse velocity in concrete. Also, aggregates in 

the concrete have the ability to scatter the pulse which significantly decreases the accuracy of 

the recorded data [Gheorghiu, Labossiére et al. 2005]. Finally, large transducers are required to 

test large concrete specimens, and the transducers are difficult to manufacture.

3.2.3.4 Fiber Optic Sensors

The use of Fiber Optic Sensors (FOSs) in civil engineering structures is becoming increasingly 

common due to the fact that the sensors are lightweight, durable, and highly accurate [ASTM

2010]. Other beneficial characteristics unique to these sensors are that they are unaffected by 

electromagnetic and radio frequency interference (which causes most traditional sensors to 

malfunction over time), no electrical power is needed at the remote location, and they have an 

excellent ability to transmit signals at long distances. Also, fiber optic sensors have a wide 

temperature range over which they can be used and do not significantly affect the stress state of 

the test subject due to their minimal size [Tozser and Elliott 2000]. 

FOSs are dielectric waveguides, cylindrical in shape, and are usually made from silica glass. 

The most common applications to civil engineering structures usually involve the sensors being 

embedded during construction. As strains develop in the structure the FOSs will also expand 

and/or contract. The light transmitted from one end of the fiber to the other by total internal 

reflection is then modulated according to the change in length of the sensor. Numerical 

measurements of the length change are taken in the form of reflected optical signals and are 

transmitted back to an analytical device. Strains within the structure are then recorded as 

indicated by the numerical measurements. See Figure 3-11 shown below.

Figure 3-11 Basic structure of a fiber optic sensor

Fairly recent studies have concluded that FOSs are capable of measuring large strains with high 

precision and without significantly affecting the bond between reinforcing steel and concrete 

[Tozser, Barker et al. 2010]. Research has also concluded that FOSs are capable of precise 

strain measurement in reinforcing steel for a variety of loading conditions, load range, and 

number of fatigue cycles [Lafhaj, Goueygou et al. 2006]. Although FOSs are most commonly 

used to measure strain, they have also be used to detect cracks within concrete structures. One 

method involves stripping the FOSs of their protective coating before embedment. When a 

crack propagates and intersects one of the sensors, the sensor is destroyed which effectively 

cancels the optical signal [Lin, Kuo et al. 2007]. The downfall of this method is that if other 

cracks propagate in the vicinity of a destroyed sensor they will not be revealed. Also, this 

method usually requires a priori knowledge of the crack region and therefore many cracks are 

missed [Lorenzi, Tisbierek et al. 2007]. Methods have since been proposed which do not require 

Core

Cladding

Primary Coating

Secondary Coating
3-54

a priori knowledge of the crack region but still require knowledge of the crack direction and so 

more sensors than is actually practical are required for accurate measurements [Phoon, Wee et 

al. 1999]. The most recent crack detecting sensors do not require a priori knowledge of the 

crack region or direction and are capable of distributed crack sensing [Nogueira and Willam 

2001]. FOSs have also been used to detect corrosion in reinforced concrete structures. One 

method uses a sensor placed around the steel reinforcement by means of super glue in order to 

measure angular strain around the bar. When the bar expands or contracts due to corrosion, the 

perimeter of its section increased or decreased and the sensor is strained [Liang and Wu 2002]. 

Although FOSs have been used to measure strain, and detect cracks and corrosion in 

reinforced concrete buildings, bridges, pavements, hydroelectric dams, tunnels, pipelines, piers, 

and piles [Tozser and Elliott 2000], it is important to note the shortcomings associated with this 

technology. First, the main application of FOSs for PCCVs would involve embedding them in the 

structure during construction. This means that all PCCVs currently in operation cannot benefit 

from this technology. Second, it has been suggested that the protective coating in these fibers 

has an effect on the strain readings they produce and may also have a slight effect on the 

response of the steel depending on the size of the instrumentation [Santhanam 2010]. Third, 

great care must be taken during the installation procedure in order to ensure that the FOS will 

perform as designed during operation. Choosing a quality adhesive and obtaining sufficient 

bond length is of particular importance. Finally, although the FOSs themselves are relatively 

cheap, the actually use of this technology in PCCV structures may prove to be an expensive 

option since they require expert installation and continual monitoring during service. Still, FOSs 

may be one of the most applicable technologies for real-time structural monitoring of future 

PCCV systems. 

3.2.3.5 Magnetic Methods

The remnant magnetism method applies a magnetic field to a prestressing tendon using an 

electromagnet at the concrete surface along the length of the tendon. Once this magnetization 

is completed, the tendon will have magnetic lines of force flowing through it. The flux is zero at 

the center of the tendon strand, and is a maximum at its outer surface. Outside of the tendon, 

the flux is not detectable. However, with the existence of a flaw (crack or break) in the tendon, 

leakage flux will occur at the location of the flaw. The detection of this leakage flux is what 

signifies damage to the tendon. This method does best identifying damage perpendicular to the 

direction of the tendon. If cracks form along the length of the tendon, they will not necessarily be 

detected [Malhotra and Carino 2004]. 

Work has been done recently to improve the state-of-the art in using magnetic fields to assess 

structural health. Of import to the nuclear industry is the improvement in the NDE technology 

which allows investigators to locate fractures in both grouted and ungrouted prestressing 

strands using the remnant magnetism method. It is possible for this method to detect fractures 

in individual strands, even if those stands are surrounded by other strands or are not detectable 

to the naked eye. This system does not require continuous monitoring of the structure, or a 

pristine condition to compare against. However, it is necessary to have access to the entire 

length of the tendon in order to be able to apply the magnetic field and test for disturbances. 

Recent experiments have been able to detect fractures in individual strands in both grouted and 

ungrouted prestressing tendons in both the laboratory and in the field on bridge beams, 

factories, indoor pools, post-tensioned bridge slabs, prestressed parking decks, and 

circumferential tendons of oil tanks. One limitation of this method is the amount of concrete 

cover through which the magnetic flux leakage can be detected. For structures with mild 

reinforcement, flaws in single strands were detectable with a concrete cover of up to 30 mm (1.2 

inches). With dense rebar, and forty strands composing the tendon, a flaw in a single strand 
3-55

could only be detected with a concrete cover of 10 mm (0.4 in.) or less [Scheel and Hillemeier 

2003]. 

There is a lot of promise to this method, but some testing of a typical PCCV tendon and rebar 

layout with typical concrete cover is recommended to assess its applicability to the nuclear 

regulatory environment. There is the potential that permanent sensors could be placed on the 

PCCV so that the structure could be tested along the same schedule as the leakage tests are 

performed, but this method also allows for the possibility of more frequent testing with 

permanent sensors in place.

3.2.3.6 Eddy Current

In Eddy Current Analysis (ECA), an alternating current (AC) is applied to what is called the 

primary coil. The AC current induces a magnetic field around the primary coil which causes a 

constant and predictable voltage in an adjacent second coil also known as the pickup coil. The 

primary and secondary coils are close but do not touch. When this system is placed near a 

conductive material, such as a metal, the primary coil’s magnetic field induces circulating 

currents within the material. These circulating currents are called Eddy currents. When a 

structural flaw begins to propagate near the coils, the Eddy current is interrupted and its 

effective impedance (the ratio of voltage amplitude to current amplitude) is increased. Changes 

in the material can therefore be detected as a change in the coil’s voltage is detected. 

ECA has been used to detect surface cracks, sub-surface cracks, and corrosion in metals, 

although it is mainly used for surface crack detection. ECA has also been used for inspecting 

the health of welds and bolt holes in metal structures [Sidek, Kabir et al. 2011]. The applications 

of ECA are mainly limited by the type of material in question and the depth of the flaw. Thin and 

highly conductive materials benefit greatly from ECA technology as the accuracy of ECA has 

been shown to decrease with non-conductive materials and with depth of the material [Deng 

and Cai 2007]. Since ECA relies on electrical current and minimal thickness to function 

accurately, its application to grouted tendons and PCCVs is limited as the strands are 

embedded deep within the non-conductive concrete.

There are several other limitations associated with ECA in PCCV applications. First, if a crack or 

corrosion area lies parallel to the current path no significant interruption will be detected and the 

crack will go undetected. Second, if electrical noise or other conductive materials are nearby the 

ECA will likely malfunction. Finally, small and insignificant geometric defects may exist in the 

structure which the ECA may term as structural flaws. Experts are therefore usually required to 

interpret the data produced by ECA which adds to the already costly application of ECA 

technology [Kenel, Nellen et al. 2005]. 

3.2.3.7 Ground Penetrating Radar

Ground Penetrating Radar (GPR) is an electromagnetic NDT method whereby a high frequency 

radio wave is emitted from an antenna into an object and returned. The antenna records 

variations in the return signal’s energy from the material which depends on the type of the 

antenna and the material being tested. After the return signal is received it is transformed into 

visual images of the sub-surface material. The interpreted images provide substantial details 

regarding the internal health of the structure [Malhotra and Carino 2004]. 

Within the last decade GPR has been used for estimation of in situ reinforcing bar size, location 

and dimensions of voids, crack detection, and corrosion detection in reinforced concrete and 

other civil engineering structures. In particular, GPR has been used for assessing the structural 
3-56

health of bridges, soils and foundations, masonry buildings, monuments, buried pipes, asphalt, 

and ducts [Rao, Kumar et al. 2007]. The extensive use of GPR in the civil engineering industry 

may allow the technology to be more easily implemented in PCCV structures. In particular, for 

degradation relatively close to the surface of the concrete member, GPR may prove to be an 

efficient way to obtain information. 

It is important to note some of the disadvantages associated with this technology. As with many 

electrically-based NDT methods, GPR is sensitive to foreign signals and will not function 

properly if unwanted electronic noise is present. Foreign signals in and around PCCV structures 

could potentially be caused by power lines, cellular phones, two-way radios, and microwave 

transmitters. Also, similar to Eddy Current, the accuracy and effectiveness of this method 

decreases as the depth of the material increases and experts are usually required to accurately 

interpret the data recorded by the device [Rao, Kumar et al. 2007]. Implementation may also be 

quite costly and may interfere with plant systems.

3.2.3.8 Impact Echo

Impact Echo (IE) technology emits mechanical stress waves through the test subject. The stress 

waves change paths as they encounter material changes in the subject. When the wave 

encounters a material defect in concrete, for example, the wave can no longer travel in a 

straight line and is reflected back to the surface. Upon reaching the surface the wave returns 

back into the concrete and proceeds to oscillate back and forth between the surface and the 

defect. The frequency at which the oscillation occurs within the material allows the depth and/or 

locations of material defects, changes, and/or voids to be determined [Sansalone and Street 

1995]. Impact echo is similar to GPR technology in that it sends a wave through the test subject 

and uses the measurement of the reflection response time to provide details regarding the 

location of internal defects. The major difference between the two technologies is that impact 

echo utilizes a mechanical stress wave while GPR utilizes an electromagnetic wave. This 

provides a slight advantage for IE technology in that mechanical waves have the ability to 

penetrate reinforcing steel while electromagnetic waves cannot do so effectively.

IE has successfully been used to locate flaws and defects in highway pavements, bridges, 

buildings, tunnels, dams, piers, piles, caissons, and sea walls. The technology has also been 

used to evaluate the early age of concrete hardness, to trace crack propagation, to detect steel 

corrosion damage in reinforcement, and to measure the thickness of concrete foundations and 

walls with an accuracy of 3% or better [Carino 2001]. It should also be noted that this 

technology is considered to be among the easiest non-destructive evaluation (NDE) methods to 

implement as the main task of the operating engineer is to simply interpret the near real-time 

frequency data to gauge defect depth. 

Since IE is mainly a defect locator, it should not be expected to perform PCCV health monitoring 

in and of itself. Although it is greatly beneficial to locate defects, the inspectors still require 

detailed information on the type of flaw propagating within the concrete, grout, and/or tendons. 

IE could be used to locate a crack in the concrete or to locate a corrosive defect in the tendons 

but not provide information on whether the defect was a crack or corrosion. Combining IE with a 

method such as radiography could prove to be very beneficial as IE could be used as the defect 

locator. Once the location of the defect is known, radiographic images could be used to provide 

details on what the defect actually is and the effect it is having on the structure. 

Several characteristics of IE make the technology somewhat impractical for use in PCCV health 

monitoring. The technology is costly to implement over large areas since most of the 

transducers are small and can only cover small surface areas. Therefore, the amount of time 

and money necessary to analyze an entire PCCV structure with IE greatly diminish its use 
3-57

independent of other NDE methods. Another disadvantage with this method, as was previously 

mentioned, is that it cannot provide details on what exactly is causing reflections to the test 

subject. Large aggregates, small voids, steel, and other minor material characteristics and/or 

flaws could generate a reflection which may not necessarily mean that a significant structural 

defect even exists [Krause, Bärmann et al. 1997]. Again, this should prompt inspectors to use 

this method jointly with other NDE methods and not singularly. 

3.2.3.9 Radiography 

Radiographic technology emits invisible electromagnetic radiation rays through a body which 

can penetrate the medium without changing direction. The density and thickness of the test 

subject will cause the emitted X or Gamma rays to gradually lose intensity [McCann and Forde 

2001]. It is through these attenuating differences that images are constructed via radiographic 

film. For example, the attenuating characteristic differences of steel, concrete, and air have 

allowed radiographic technology to locate the reinforcing rebar within bridges.

Radiography has been used extensively in the civil engineering industry, particularly for 

assessing the health of reinforced concrete structures. The technology has been successful in 

locating internal cracks, voids, and variations in the density of concrete. The technology has 

also proven successful for assessing the state of both grouted and ungrouted tendons in 

prestressed box girders [Bligh, Nakirekanti et al. 1995, Owen 1998]. Uniformity of the cement 

grout, the condition of the cable sheaths, and strand health have all been monitored with 

Radiography which makes this technology directly applicable to PCCV health monitoring. The 

location and health of the tendons and the state of the grout and concrete, and the bond 

between the grout and tendons could all be assessed with radiography. 

Although Radiography has been successful in many civil engineering venues, there are several 

limitations associated with the technology. The first is that the users are exposed to radiation 

throughout the duration of each test which poses an obvious and significant safety hazard to the 

users. Still, the issue may not be as significant in PCCVs as it would be elsewhere because 

plant operators are trained to handle the hazards associated with radiation [Mehrabi 2006]. 

Second, the technology is fairly expensive and time-consuming to implement. Finally, energy 

requirements for the technology can be quite demanding and the output images are not always 

easy to interpret. As with many of the other NDT technologies previously discussed, experts are 

usually required to interpret the visual results.

3.2.3.10 Vibrating Wire Strain Gauges

Vibrating wire strain gauges (VWSGs) utilize electromagnetic and vibration technology to 

provide information on the health of a structure exposed to both static and dynamic loads. A 

typical VWSG is composed of two round flat ends joined by a protected steel wire. Deformations 

in the test subject are transferred to the flat ends of the gauge and induce variations in the 

length and vibration frequency of the gauge wire. The change in length provides strain data to 

the user and is monitored by an electromagnet which sits in the middle of the gauge. Many 

VWSGs also have built-in resistive temperature sensors and are therefore able to monitor 

temperature fluctuations [Barr, Stanton et al. 2005]. The general structure of a VWSG is shown 

below in Figure 3-12. 
3-58

Figure 3-12 Depiction of a Vibrating Wire Strain Gauge

The main applications of VWSGs consist of either embedding them during construction or 

attaching them to the surface of the structure. Successful VWSG applications to date include 

but are not limited to monitoring strain, temperature, and pressure changes in dams, bridges, 

viaducts, large buildings, tunnel linings, piles, and pressure vessels. Therefore, the strains, 

temperature changes, and pressure changes in the concrete, grout, and tendons of PCCV 

structures have the potential to be monitored with this technology. Also, since recent studies 

have shown that embedded VWSGs can provide accurate data for decades; this technology has 

potential to monitor PCCV structures over the course of its entire life [Frangopol, Strauss et al. 

2008]. 

Present day PCCVs cannot fully benefit from VSGW technology because embedded gauges 

must be cast in place during construction. Although it is possible to measure surface strains and 

temperature changes in currently operational PCCVs, internal damage would go undetected. It 

should also be noted that these gauges cannot reproduce variations in concrete deformations 

accurately until the concrete/grout in which it is cast has time to cure to a sufficient stiffness. 

Only when the concrete has reached sufficient stiffness is the gauge considered to be perfectly 

attached and therefore transmitting accurate strain measurements. Also, if the gauges are used 

to monitor grouted tendons it is imperative that the gauges be expertly attached to the tendons 

during casting. Other disadvantages associated with this technology are that they tend to be 

relatively large, they require bulky cables for power supply and signal transmission, and are 

fairly expensive. 

3.2.4 Limitations of Monitoring for Grouted and Ungrouted Tendon Systems in 

PCCVs

For the basic techniques of monitoring (i.e., those employed in the nuclear and the general 

building industry), monitoring systems are generally the same between grouted and ungrouted 

tendons. They consist of:

• Measuring tendon forces (Pjack) during installation

• Measuring tendon elongations during installation

• Measuring structure deflections as a function of prestressing or pressurization

For non-nuclear structures, there generally are no monitoring programs employed during the 

long-term life of the structure.
3-59

For ungrouted tendons, especially nuclear structures, there are more possibilities for monitoring:

• Load cells left in place on selected tendons

• Strain-gauge (or other strain measuring device) installed on tendons; technology has 

advanced substantially in recent years which could greatly facilitate such measurement, 

especially through the use of wireless measurement, where scans of the measurements 

can be taken from outside the structure.

• Sound-monitoring of tendon systems – to listen for redistribution of strands, and in 

extreme situations, individual wire-breaks. Such individual wire-breaks have often been 

found to serve as a pre-cursor of a tendon overstress well before an entire tendon fails

• Other NDT methods mentioned above.

With grouted tendons, of utmost difficulty is detecting corrosion in tendons and damage to 

tendons prior to failure of the tendons. Even if ungrouted tendons are left to assess the state of 

prestressing over the life of the structure, these do not represent the state of grouted tendons 

which potentially have corrosion. 

The state of the art in terms of NDT methods should considered for use in monitoring PCCVs 

with grouted tendons. Various countries are exploring the usefulness of such methods. Because 

many of these methods are still in their infancy, regulatory requirements cannot necessarily be 

made. The use of these technologies provides the opportunity to increase the overall safety and 

confidence in the health monitoring of plants. 

Table 3-7 Summary of NDT methods and their applicability to monitoring prestressed 

PCCVs.

Useful for nuclear applications 

(where used / most experience)

Potential for nuclear 

applications

Not as useful for 

nuclear applications

Visual (Around World) 

Acoustic (Canada)

Ultrasonic Pulse Echo (US & World)

Fiber Optics (Finland, Canada)

Magnetic Methods

Impact Echo

Radiography

Vibrating Wire Strain Gauges

Ground Penetrating Radar 

Eddy Current

3.2.5 Corrective Measures in PCCVs

The repair and replacement of existing ungrouted tendons is covered in the 2010 ASME Boiler 

and Pressure Vessel Code XI Rules for Inservice Inspection of Nuclear Power Plant 

Components [ASME 2010b]. However, there is no set practice for grouted tendons. 

There is no easy way to remove grouted tendons for repair due to the grout along the entire 

length of the tendon. If local corrosion is identified, such questions arise as to whether the

tendon needs to be replaced, and at what point does damage need to be repaired. The changes 

in overall response to local tendon damage are investigated in Section 4.10.

Another potential issue with grouted tendons is the difficulty in repairing other parts of the 

nuclear power plant, for instance replacing a steam generator. With ungrouted tendons, the 

tendons can be removed from a section of the wall which can then be removed and repaired. 
3-60

This is not as simple in grouted tendon systems. The tendons cannot be de-tensioned and 

removed easily. 

3.2.6 Summary and Conclusions

Throughout the world, the regulation of grouted tendons for use in nuclear power plants is still 

being developed. Much of the world has relied on the US NRC’s regulatory guidance on many 

regulatory issues, for use in their own countries. Other countries such as Belgium, Canada, and 

France have more operating plants with grouted tendon systems and have developed their own 

regulatory practices. This document serves to consolidate this (and other) experience. Some 

plants with grouted tendons have seen many years of service without reported corrosion 

damage of note to date in the grouted tendons, however corrosion is difficult to measure directly 

for grouted tendon systems. The effect of corrosion may be indirectly monitored through 

instrumented measurement of containment deformation. 

Despite the lack of reported corrosion damage in PCCVs with grouted tendons, there have been 

sufficient corrosion problems in the non-nuclear industries using grouted tendons to warrant the 

need for improved monitoring methods of grouted tendons. Because the prestressing level in 

grouted tendons cannot be measured directly, the use of non-destructive testing methods

should be considered and developed for use in nuclear power plants. Various NDE methods are 

presently being implemented and studied by countries and regulatory bodies around the world. 

With the new construction of new containments in Europe that incorporate grouted tendons, 

valuable experience and information will be gained.
4-1

4 ASSESSMENT OF DURABILITY AND LONG-TERM CORROSION 

PROTECTION

The focus of the third section is the assessment of long-term corrosion protection methods to 

ensure life-long durability for the grouted and ungrouted tendon systems currently used in 

nuclear power or related industries. A review and discussion of information and / or data related 

to corrosion that concerns tensioning element strength and ductility, tendon sheathing filler 

characteristics and tendon duct type is presented. Other aspects (e.g., aging mechanisms, 

stress) of material physical condition were addressed to assess durability.

4.1 Introduction

This section focuses on a comparison of post-tensioning methods and inservice inspection 

requirements for grouted and ungrouted tendons. The first few sections discuss the corrosion 

process for post-tensioning methods, with comparisons for grouted versus ungrouted tendon 

systems. Of interest is the difference in methodology for installing these systems, and ensuring 

that such differences in post-tensioning methods are not detrimental to structural health in long 

term behavior. Section 4.10 documents work on parameter sensitivity studies using FEA to 

model postulated tendon corrosion, and examine the predicted effects on structural 

performance. Section 4.12 discusses the monitoring methods used to assess the structural 

health of grouted and ungrouted tendon systems, and compare practices in a variety of 

industries which use the systems. 

The basis for the FEA, and for much of the general discussion of PCCV behavior is the 1:4 

Scale PCCV Model Test and Sandia’s extensive containment research program, dating back to 

the 1980s. Both general subjects are summarized in [NUREG/CR-6809 2003, NUREG/CR-6810 

2003, NEA/CSNI/R(2005)5 2005, NUREG/CR-6906 2006]. 

Most prestressed containments in the USA were constructed with ungrouted tendons to simplify 

inservice inspection and testing. One operating US PWRs has a prestressed concrete 

containment with grouted tendons: Robinson Unit 2 has grouted, vertical-bar tendons. 

Additionally, Three Mile Island Unit 2 (no longer in operation) has grouted strand tendons. 

Grouted tendons have been used in 34 900-MWe French PWR containments [Barbe and 

Costaz 1991]. In containments with grouted tendons, the concrete is poured around metal ducts 

in which the tendons are inserted. Once the concrete has cured, the tendons are tensioned and 

the ducts are filled with grout. After the curing of grout, the external tension in the tendons, 

which is applied with hydraulic jacks, is released. The bond now developed between the grouted 

tendon and the concrete places the concrete in compression.

The most serious threat to the durability of any prestressed concrete structure is corrosion of the 

prestressing and/or reinforcement [Collins and Mitchell 1997]. Rebar corrosion can cause 

spalling of cover concrete (due to corrosion induced expansion and loss of bond). Such damage 

is very difficult and expensive to repair. Corrosion of prestressing tendons is of even greater 

concern, because it can trigger collapse of a structure, or in less severe corrosion conditions, 

loss of functionality for which the structure was designed. As described previously, in the history 

of prestressed structures, there are many examples of bridges experiencing problems, and 

similar issues can be cited for buildings. 

In general, nuclear power plant concrete structure’s performance has been very good with the 

majority of known problems occurring during construction or shortly after construction (and 

corrected at that time). However, with passing time, aging of concrete structures occurs, and if 
4-2

potential degradation effects are not managed, an increase in the risk to public safety can occur. 

There are three main degradation mechanisms that can cause aging damage to concrete 

containments, even after a long period of satisfactory operation: 

1. alkali–silica reactions (ASR); 

2. corrosion (in various forms) of reinforcing steel, steel liner (especially the embedded portion 

of the liner) and prestressing tendons; and 

3. sulfate attack (particularly magnesium sulfate attack that causes loss of concrete strength). 

Initially, the occurrence of these mechanisms is not readily observable, however, as the 

degradation progresses, there is the potential to cause widespread concrete cracking and 

strength loss. ASR has been detected at the Seabrook plant in the US in several concrete 

structures including the reactor containment building. Leaching and sulfate attack, which 

produces expansive products, occurs after a relatively short exposure and is detectable at the 

concrete surface. Such damage has not been reported in containment vessels, but it has been 

reported in many non-nuclear structures. Long term prestressing losses have also been an 

issue - rates at some US plants have exceeded the design limits; the prestress losses that were 

predicted to occur after 40 years of operation were detected during the 20 and 25-year tendon 

inspections. This issue of loss of prestress can be addressed, if necessary, by re-tensioning 

(if ungrouted tendon systems are used), but it is difficult or impossible to address with grouted 

tendons. 

Because of experiences with corrosion in tendon systems in general (bridges, buildings, and 

other structures), corrosion protection remains an important consideration for PCCV design, 

maintenance, and long-term performance.

4.2 Chemistry of Corrosion Process

A number of texts on prestressing (e.g., [Collins and Mitchell 1997]) provide useful information 

about corrosion of prestressing and rebar; we summarize the process here.

Corrosion of steel inside of concrete is an electrochemical process (see Figure 4-1) similar to 

the process that takes place in a battery. The process is similar whether referring to rebar or to 

post tensioning tendons. (What differs between rebar and tendon steel is the speed of the 

corrosion and the amount of section loss – post tensioning tendons tend to be even more 

vulnerable than rebar due to its chemical composition.) One part of the steel acts as the anode. 

This is where the steel corrodes; the iron is oxidized to ferrous ions and electrons are given off. 

Other areas of the steel, which have a higher electrochemical potential, act as cathodes, 

consuming oxygen, water, and electrons to form hydroxyl ions. To complete the electrochemical 

cell, the bar or tendon segment acts as the electrical conductor and the concrete pore water 

containing dissolved salts acts as the electrolyte. 
4-3

Figure 4-1 Corrosion of reinforcement – an electrochemical process. Adapted from

[Collins and Mitchell 1997].

Portland cement concrete is typically highly alkaline owing to the mainly to the large number of 

calcium ions produced in the dissolution of Portland cement. Even after the cement sets, the 

hydration of the Portland cement produces large quantities of calcium hydroxide, which makes 

the system stay at a pH level of about 12.5. For the range of typical potentials for concrete and 

at this level of alkalinity, the steel will typically be passive. The passivity of the steel is 

characterized by a thin and tightly adherent oxide film on the surface of the steel which protects 

the steel from further corrosion. This is similar to the oxide films that form on the surface of 

aluminum or zinc under normal atmospheric conditions and prevent further corrosion. While the 

oxide film remains unbroken, the steel will not further corrode. 

The natural alkalinity of concrete can be reduced by the chemical reactions that occur between 

the hardened Portland cement paste and carbon dioxide. This process, which is called 

carbonation [Papadakis, Vayenas et al. 1989, Papadakis, Vayenas et al. 1991, Papadakis, 

Fardis et al. 1992a, Papadakis, Fardis et al. 1992b], can lower the pH of the concrete to about 

8, destroying the passivity of the embedded reinforcement and permitting corrosion to start.

Cathodic

Area

Cathodic

Area

Anodic 

Area

Electron 

Flow

Environment: H2O, O2, Cl-

, CO2

Fe OH ++ OH- -

Crack

Reduction Oxidation Reduction

Concrete or 

Grout

Rebar or Tendon

Rebar or Tendon
4-4

Fortunately, carbonation is a very slow process, particularly in high-quality concretes. 

The presence of chlorides can endanger the passivity of the steel even at high pH levels. Very 

small concentrations of chlorides present when concrete is in the liquid phase can destroy the 

protective oxide film on the steel. The chlorides in concrete come from many sources. Typically, 

the cement itself will contain about 49.9 to 99.89 mg/L (50 to 100 ppm) by weight of chlorides. 

Potable water usually has less than 249.7 mg/L (250 ppm), with city water supplies generally 

having less than 49.9 mg/L (50 ppm). In comparison, seawater contains about 19,977 mg/L 

(20,000 ppm). The chloride content of the aggregates generally fall in the range of 9.99 to 399.5 

mg/L (10 to 400 ppm). Some dolomites can contain 998.9 mg/L (1000 ppm). Some admixtures 

also tend to add chlorides.

More important than the chlorides in the mix ingredients are the chlorides that may penetrate 

the concrete during its service life. Structures in or near seawater will be subjected to saltwater 

or windborne spray. In regions that use deicing chemicals, parking garages, bridges, and 

structures near roads will all be subjected to highly concentrated chloride solutions.

The ACI Code, the CEB-FIP Code, and other concrete codes contain detailed information and 

requirements about concrete cover requirements and about allowable crack widths for concrete 

in service. For example, the CEB-FIP [CEB-FIP 1978] allowable crack recommendations are 

shown below in Table 4-1.
4-5

Table 4-1 Allowable concrete crack widths from [CEB-FIP 1978].

Exposure conditions

Load to be 

considered

Limits for 

reinforcement highly 

sensitive to 

corrosion

Limits for 

reinforcement 

moderately sensitive 

to corrosion

Mild

Usual exposure

Low-humidity exterior 

exposure

Frequent (dead load 

plus frequently 

occurring live load)

0.2  0.3 mm

(0.008  0.012 in)

0.4  0.6 mm

(0.016  0.024 in)

Permanent (dead load 

plus sustained live 

load)

0.1  0.15 mm

(0.004  0.006 in) Need not be checked

Moderate

High-humidity or 

slightly corrosive 

interior exposure

Running water

Ordinary soil 

exposure

Usual exterior 

Exposure

Frequent 0.1  0.15 mm

(0.004  0.006 in)

0.2  0.3 mm

(0.008  0.012 in)

Permanent No Tension in 

Concrete Need not be checked

Severe

Seawater exposure

Slightly acidic liquids

Deicing Chemicals

Corrosive Gasses

Corrosive Soils

Rare (dead load plus 

maximum possible 

live load)

0.1  0.15 mm

(0.004  0.006 in)

0.2  0.3 mm

(0.008  0.012 in)

Frequent No Tension in 

Concrete

0.1  0.15 mm

(0.004  0.006 in)

Cracks in PCCVs subjected only to routine operation loads ought to be rare or non-existent, 

because they are prestressed structures. But it is possible for small cracks to exist near 

anchorages due to the microcracking that occurs during jacking. And since this is where 

prestressing systems are near the surface of the concrete, this can make anchorages 

particularly vulnerable to corrosion.

While corrosion of the steel is the greatest threat to the durability of prestressed concrete 

structures, other types of chemical attack can also occur. For example, solutions containing 

sulfates attack concrete made from normal Portland cement, and certain types of aggregates 

chemically react in the alkaline environment of the concrete.

4.3 Corrosion Over Time

4.3.1 Aging Mechanisms Related to Corrosion

As described in the preceding subsection, carbonation is a very slow process, particularly in 

high-quality concretes. Thus for a concrete where fc’ = 35 MPa (5,000 psi), the depth of 

carbonation would probably be less than 10 mm (0.4 in) after 20 years. Hence, in uncracked 

concrete, carbonation will not typically penetrate the concrete cover. If the concrete contains 
4-6

cracks, the carbon dioxide can reach further into the concrete and the depth of carbonation will 

be increased cracks could arise from thermal or drying shrinkage, alkali silica reactivity, or 

freeze-thaw cycling.

The presence of chlorides can endanger the passivity of the steel even at high pH levels. Very 

small concentrations of chlorides in the liquid phase of the concrete can destroy the protective 

oxide film on the steel. The chlorides in concrete come from many sources. 

Once the passivity of the embedded steel has been destroyed, either by carbonation of the 

concrete or penetration of chlorides, corrosion will start and the remaining life of the structure 

will then depend on how fast the corrosion proceeds. The rate of corrosion depends on how 

efficiently the electrochemical cell can operate. In most situations, it will be the electrical 

resistivity of the concrete and the availability of oxygen at the cathode that will control the rate of 

corrosion. If the electrical resistivity of the concrete is high enough, not enough current can flow 

to generate significant corrosion. While the quality of the concrete will influence the resistivity, it 

is the moisture content of the concrete that is the most dominant parameter determining 

resistivity. If the degree of water saturation is less than about 40%, the resistivity will be high 

enough to prevent significant corrosion. 

A secondary factor in the corrosion process for steel embedded in concrete or grout is the 

availability of oxygen. The permeability of the concrete and the thickness of the concrete cover 

over the steel are the most important factors governing the availability of oxygen. 

4.3.2 Comparison of General Lifetime of Tendon Systems as a Function of 

Corrosion

Specific field experience over time with U.S. nuclear power plants is cited by Shah and 

Hookham [Shah and Hookham 1998]. A limited number of corrosion-related failures of 

containment tendons and other prestressed structure tendons have been reported in both 

nuclear and nonnuclear industries (parking structures, bridges, storage tanks, buildings). Many 

of these problems were the result of poor construction practices such as inadequate protection 

of end anchorage components, inadequate concrete cover, inadequate grease or grout 

coverage, contaminated grease, electrical contact with other embedded steel, voids inside or 

outside ducts, and the presence of water inside ducts. The literature review conducted for the 

current work did not uncover quantitative corrosion versus time data for corrosion of 

prestressing in nuclear power plants.

A review of ungrouted tendon surveillance reports indicates that corrosion has been detected at 

a number of plants during both pre-service, and inservice inspections. The most common 

incidents of corrosion are pitting of the tendon wires or strands. The pitting in most cases has 

been limited and attributed to the presence of small amounts of water that have accumulated at 

the lower end of the vertical tendons in the area of the anchorage. This problem has been 

encountered in both Types I and II post-tensioned PWR containments where clogged or blocked 

drains cause rain water to collect on the dome and seep through the top tendon anchorage and 

travel to the bottom anchorage. Another reason for pitting corrosion of the tendon wires and 

anchor heads is poor construction practice that resulted in the tendons being stored at the site 

for long periods without proper protection, or not being properly protected after installation and 

before application of the permanent corrosion protection material. A few anchor heads on some 

containments have experienced either cracking or partial failure and, in one case, a complete 

failure because of hydrogen embrittlement. The time during which these failures occurred 

ranged from a few days to a few years after the tendons were tensioned. The chemical 

composition of the anchor head material and the heat treatment procedures used are thought to 

have contributed to these failures.
4-7

Three anchor heads at the bottom of the vertical tendons at Farley Unit 2 were found in a failed 

condition (two broken and one cracked) after 8 years of plant operation [Naus, Marchbanks et 

al. 1988]. Water was found in the grease cap or under the anchor head in each case. The 

quantity of water varied from a few ounces to one-half pint. Metallurgical evaluation of the failed 

anchor heads identified hydrogen-induced cracking as the failure mechanism. High-strength 

steels, such as anchor-head steel, subject to sustained high tensile stresses are susceptible to 

this mechanism if a source of atomic hydrogen is present. 

General corrosion and stress corrosion cracking of tendons have been observed at the 

prestressed concrete reactor pressure vessel for the Fort St. Vrain plant, which was a gascooled reactor that is now shut down. Several instances of corroded wires and broken tendons 

were discovered during the 1984 tendon surveillance. The corrosion was caused by the 

products of microbial attack on the anticorrosion grease in the tendon sheaths [NUREG/CR4652 1986]. The licensee had proposed to halt the corrosion damage by filling the tendon 

sheaths with inert nitrogen and increasing the frequency of visual inspection and lift-off tests.

Several failures of the prestressing steel have been reported in nonnuclear applications. 

Prestressed concrete tanks have failed as a result of corrosion, stress corrosion cracking, and 

hydrogen embrittlement of reinforcing steel. The Berlin Congress Hall, built in 1957, collapsed in 

1980 due to the hydrogen-induced stress corrosion cracking of grouted tendons (so this was 

after 23 years of service) [Field and Carper 1997]. The failure was the result of a construction 

error; the failed tendons were not sufficiently surrounded by grout. It appears that hydrogen 

produced by a cathodic reaction might have contributed to this failure.

Some of the best tendon corrosion versus time data available was tested and recorded by the 

University of Texas on Segmental Bridge research for the U.S. Federal Highway Administration 

and Texas Department of Transportation [Research Report 1405-6 2002]. The report provides 

formulae for corrosion as a function of time, and of corrosion current density. For various 

corrosion conditions (caused by various chloride exposures, grouts, and duct types), corrosion 

assessments were made at 580 days, 640 days, 710 days, 1250 days, 2347 days, 2712 days, 

and 2782 days (8 years). Corrosion ratings were assessed for corrosion and pitting on individual 

wires of 7-wire strands. Under the most severe conditions, all exterior wires (6 out of 7) were 

highly corroded. Translating this into percent section loss requires some engineering judgment, 

but based on the research, it is possible to envision tendon section losses ranging from just a 

few percent (a wire or two breaking) to 70 or 80% of the section.

4.3.3 Progression Characteristics of Corrosion in Each Type of System

Corrosion initiation for prestressing systems is most likely near the anchorages. Once it begins, 

it is important to have information about the mechanism and rate for spread of corrosion. It is 

also possible for moisture and oxygen to reach an interior location of a tendon if there is a ductjoint failure (and these are often associated with a grout air pocket), so the same questions exist 

if the corrosion were to begin at a location far from an anchorage. 

These questions are addressed in some detail by the ACI Committee Report 222R-01 [ACI

2001]. The following is a summary of the information provided. The report also contains an array 

of useful references.
4-8

Most literature discusses Faraday’s Law when it comes to the rate of corrosion in any metal

[ACI 2001]. Corrosion rate can be determined as a current by measuring rate at which electrons 

are removed from the iron in the anodic reactions. Faraday’s Law is as follows:

w ItA M

nF = (4-1)

Where M is mass of metal dissolved or converted to oxide, t is the amount of time in seconds 

the metal has been allowed to convert to its oxide and/or dissolve, I is the current in Amps, Aw 

is the atomic weight, n is the valency, and F is Faraday’s constant. 

Active corrosion can only proceed when the protective passive film begins to break down. This 

is because when the film breaks down an increased supply of oxygen is allowed to come into 

contact with the metal and increases the cathodic current. This significantly increases the 

corrosion rate. The film can break down over the entire metal due to significant changes in the 

general thermodynamic conditions which is generally due to a significant decrease in pH which 

leads to an unstable passive film. The film can also break down locally by way of a chemical 

attack, e.g. chlorides, or mechanical failure such as a crack in the concrete cover. 

The actual detailed mechanism of breakdown of the passive film by chlorides is not known 

because of the difficulties associated with examining the process on an atomic scale in the 

extremely thin passive layers. Some theories are available. However, it is known that if chlorides 

are used as admixtures, the electrical conductivity of the system is increased which will act 

adversely with the steel to increase the corrosion rate. Active chloride corrosion generally 

produces a light green semisolid reaction product near the steel, which if exposed to air, will turn 

black and begin to rust. Since the iron hydroxides have a larger specific volume than the steel, 

internal stresses are developed within the concrete which may cause cracking and/or spalling. 

The expansive effect of the corroding steel is not considered in the analyses presented here. 

The effect of chlorides on the system may not always be negative but is a complex function of 

several parameters including C3A (Tricalcium Aluminate), C4AF (Tetracalcium Aluminoferrite), 

pH, w/cm, and whether the chlorides were added into the initial mixture or penetrated the cured 

concrete. Some research has also shown that the Cl/OH ratio is also important and should be 

low (< 0.3) to keep corrosion from initiating.

Carbonation has also been shown to be an initiator of steel corrosion. Carbonation generally 

proceeds in concrete as a front. Upon reaching the steel, depassivation over large areas of the 

steel can occur which can easily cause corrosion to initiate. Carbonation-induced corrosion is 

influenced by thin concrete cover, cracks, and high porosity coupled with a high w/cm.

Depassivation is necessary but not sufficient for corrosion to continue to propagate. Moisture 

and oxygen are essential if the corrosion is to proceed at a significant rate. Although chlorides 

and/or carbonation may be directly responsible for the initiation of corrosion, they do not appear 

to play a direct role in the rate of corrosion after initiation. The most recent research suggests 

that the main rate-controlling factors are the availability of oxygen, the electrical resistivity, the 

relative humidity, the pH and the temperature. 

The corrosion rate after initiation is also a function of the anode to cathode area ratio. For 

example, if a tendon somehow comes into contact with exposed steel, the exposed steel 

becomes anodic and the tendon becomes cathodic. Because the amount of embedded steel is 

often far greater than the exposed steel, the rate of corrosion of the exposed steel can be 

extremely high. In other words, if the cathode to anode ratio is much greater than 1, the rate of 

corrosion will increase significantly. Although Faraday’s law is appropriate for estimating tendon 
4-9

corrosion over time, the rate of corrosion is difficult to know with certainty. The rate of corrosion 

will differ along the length of the tendons and/or bars b/c the passivation provided by the 

concrete varies. Corrosion loss can therefore concentrate in just one portion of the strands. It is 

possible that the tendon be anodic and cathodic simultaneously in different sections of the 

strand. Faraday’s Law only provides an estimate.

4.4 Description of Long-Term Corrosion Protection Methods

A good deal of information about the tendon corrosion protection problem, and about modern 

corrosion protection methods is available from:

1) Reports and required specifications developed by the regulators of other structures (for 

example, bridges), and 

2) reports by furnishers of tendon systems, such as VSL, Freyssinet, or Dywidag. 

The following provides a summary of information presented in even more detail in References

[CEB-FIP 1990, VSL 1992, Sansalone and Street 1995, Corrosion 1996, Technical 1996, Ganz 

1997, Research Report 1405-4 1999, Interim 2000, FIB-7 2000, Matt, Hunkeler et al. 2000, 

Cullington, MacNeil et al. 2001, Matt 2001, Woodward 2001, VSL 2002, Guide 2003, Hewlett 

2004, PTI 2012].

Corrosion of prestressing system and reinforcement is usually not the root cause of a durability 

problem but rather a consequence of inadequate consideration for durability in the overall 

design of the structure. Multiple layers of protection are required against corrosion:

1) First and most important layer of protection is the overall concept and design of the 

structure. A key element in this design is to keep water off the structure and the 

reinforcement, and/or to assure that it drains quickly from the structure. 

2) A second layer of protection can be provided with water-proofing membranes in particular 

on critical surfaces exposed to water and other aggressive media such as deicing salts 

(Typical in transportation structures but not often seen for nuclear structures). 

3) A third layer of protection in concrete structures is provided with dense concrete designed 

specifically for low permeability. 

4) A fourth layer of protection for the tendons of post-tensioned structures includes the duct 

system, either steel or plastic which acts as a barrier against moisture and ion ingress. This 

layer may also include a plastic sheath around the tendon itself, independent of the duct. 

These two layers are consolidated into one logical layer here. An example of this technology 

is the Fressynet C-Range system with sheathed tendons [Freyssinet 2010]. 
4-10

5) The last layer of protection of post-tensioned structures is provided directly onto the 

prestressing steel itself in various forms. Grease or cementitious grout can each be applied 

depending on the design of the system. Grease or grout, as reviewed in detail in this report, 

is the last, and is only one, of the layers of protection of tendons in post tensioned 

structures. While high quality grease or grout placement is important for the durability of 

tendons, it alone cannot guarantee the durability of tendons. It is the owner’s and the 

engineer’s obligation to select and specify a suitable combination of independent layers of 

protection adapted to the particular environment in which the structure is built. Additional 

layers of protection provided during construction have a relatively insignificant cost 

compared with repair of durability problems of a structure in operation.

Out of 14 bridge projects with grouting defects, the majority of corrosion problems in the 

tendons were caused by ingress of water containing chlorides. A durable and leak tight 

encapsulation of the tendons, e.g. with robust plastic ducts, was considered essential to 

improve the protection and to assure the durability of grouted posttensioning tendons.

More recently in the bridge industry, durability problems due to incomplete grouting and 

corrosion have been reported in the USA - one completely and one partly failed external tendon 

were found during a detailed inspection, [Interim 2000]. During inspection many end anchorages 

of the external tendons located at the high point of the tendon profile were found incompletely 

filled with grout.

Among the most comprehensive studies of long-term aging effects and aging performance for 

LWR concrete containments (both reinforced concrete containment vessels and PCCVs) is the 

work of Shah and Hookham of INEL. The work is summarized in [Shah and Hookham 1998] and 

the findings of the work, especially those pertinent to PCCVs are paraphrased below. The 

findings, though relevant to containments, do not necessarily constitute documented LWR 

containment problems.

1) Alkali-silica reactions, corrosion of rebar, liner, and prestressing, and magnesium sulfate 

attack have the potential to cause widespread aging damage to PCCVs even after years of 

satisfactory service.

2) The threshold value of chlorides needed for corrosion initiation will be lower if the hydroxyl 

ions are reduced by carbonation, leaching, or magnesium sulfate attack. 

3) The main concern related to corrosion of prestressing is the loss of structural capacity of the 

concrete structure.

4) Grease surrounding ungrouted tendons adequately protects tendons, provided that the 

water content and the water soluble chlorides, sulfides, and nitrates contents in the grease 

are kept below the specified limits, and sufficient reserve alkalinity is maintained. Grout 

surrounding the grouted tendons adequately protects the tendons, provided there are no 

voids in the grout and the tendons are fully covered with grout.

5) Maintenance of existing surface coatings and early repair of cracks in the cover concrete are 

effective to mitigate aging damage in containment vessels.
4-11

4.5 How Corrosion Affects Tendon Element Strength for Different 

Systems and Protection Methods

In subsection 4.8, the authors investigate this subject in some detail through a series of FEA 

simulations, under the general topic of “Stress State as a Function of Corrosion”. But some 

aspects of grouted versus ungrouted structure component behavior have been studied with 

experimental work.

A good example of such studies is found from the late 1970s [ORNL/TR 6478 and ORNL/TR 

6479], when some of the currently operating U.S. nuclear fleet was still being designed, Oak 

Ridge National Laboratory performed studies of post-tensioning systems for PCCVs, and 

published a similar list of pros and cons, but the list covers a few additional items particularly 

germane to NPPs.

Flexure beams were created and tested under static and dynamic loads for ungrouted and 

grouted tendons. Different grout types were also investigated for beam strength as well as bond 

transfer length and pull-out tests. Test set-ups are described in the reference. But the 

conclusions of the study are paraphrased below:

1) The grouted tendon beam elements provide improved crack control (more cracks but much 

narrower crack widths), improved ultimate load capacities in flexure, and conservatism in 

seating and overall anchorage efficiency relative to the ungrouted tendon beam elements. 

(It should be noted that anchorage failure is more critical for an ungrouted tendon than for a 

grouted tendon.)

2) The shrinkage-compensating cement grout material produced flexure members with slightly 

improved (<10%) ultimate load capacities, prestressing bond transfer lengths 47% less, and 

bond pull-out values equivalent to those of specimens grouted with conventional materials. 

Testing was limited; however, performance improvement trends were not significant enough 

to merit a recommendation for its use over conventional grouts in PCCVs as a general grout 

material.

3) Polymer-silica cement grout material produced flexure members with a slight reduction 

(<8%) in ultimate load capacities, prestressing bond transfer lengths 61% less, and bond 

pull-out values superior to those of specimens grouted with conventional materials. Results 

indicate that polymer-silica-based grouts have application where improved bond strengths 

are required (such as for anchorages) but are not presently recommended as a general 

PCCV grout because of their relatively high costs. Elevated-temperature test results indicate 

that these materials also have application in regions of elevated temperature [Khayat 1998, 

Schokker, Koester et al. 1999], though limited information is available for combined elevated 

temperature and radiation. 

4) The flexural members fabricated from fibrous concrete demonstrated improved ductility and 

resistance to cracking relative to the conventional concrete prestressed members. These 

materials have potential application in areas of stress concentration such as at penetrations 

to reduce reinforcing steel requirements or in regions requiring improved impact resistance.
4-12

4.6 Tendon Sheathing Filler Characteristics and Effect on Corrosion 

Behavior

This subsection discusses effects on corrosion behavior of the tendon grouting systems, and 

any improvements obtainable via changes to grouting system or chemical composition. While 

grout has many interesting characteristics and qualities relating to corrosion protection, greased 

tendon systems are also quite capable of inhibiting corrosion. The following discussion is 

focused on grouted systems, but should not suggest that grout is the only method of controlling 

tendon corrosion.

4.6.1 Cementitious Grout Characteristics

The following information is summarized from a conversation with Professor George Hearn of 

the University of Colorado, Boulder, who is an expert in prestressed concrete design and 

construction.

It is safe to conclude that grouts differ mainly by additives. Non-shrink cementitious grouts are 

preferred. Shrinking grouts (used mainly in bridges before 1980) caused radial cracks in the 

grout which allowed water and air to reach the strands, thereby propagating corrosion. Cracks 

alone are not sufficient for tendon corrosion. The cracks have to occur where water is present or 

at least where water is allowed to penetrate the crack. The grout is therefore not the main culprit 

of corrosion, but instead the cracks which appear in grout due to shrinkage are the main culprit. 

After such shrinkage occurs, protection to the tendons is lost, and generally speaking, the 

tendons become nearly as vulnerable to corrosion as ungrouted tendons, or perhaps more, 

since ungrouted tendons do have grease or oil protection.

The Oak Ridge National Laboratory work cited earlier [ORNL/TR 6478 and ORNL/TR 6479], 

also addresses the issues of filler characteristics (grout composition). The following provides a 

summary of those findings. 

For the length of exposures investigated, the corrosion-inhibiting capability of grout for 

protecting prestressing steel materials has been demonstrated to be at least equivalent to that 

of commercial organic-based products in the presence of S- and Cl- environments. The 

corrosion inhibiting capabilities of grout and commercial organic-based products in NO3

-

environments were equivalent for exposures of up to 38 days; however, for greater exposure 

periods, the NO3

- environment produced ductility reductions with no decrease in load capacity 

for the grout protected specimens. It should be noted, however, that ammonium nitrate solution 

was chosen as a worst case because it readily attacks both the grout and the prestressing steel.

It is interesting to note that 50.8-mm (2 in) grout cubes placed in the NO3

- solution at 66°C (140 

F) exhibited strength decreases of 37 and 57% relative to control specimens cured in limewater 

for exposure periods of 35 and 60 days, respectively. For an exposure of 101 days, the 

specimens deteriorated to the point that they could not be tested. Similar specimens placed in 

the hydrogen sulfide and chloride solutions did not exhibit significant strength changes for 

exposure times up to 100 days.

4.6.2 Grease Characteristics

Grease used in post tensioning applications provides corrosion protection to the steel tendons 

by physically coating the tendon surface thus excluding moisture, oxygen, and deleterious ions 
4-13

from the surface of the steel. The chemistry of the grease may take a variety of formulations, but 

is generally highly viscous and hydrophobic so as not to be easily stripped from the tendon 

surface in the event of moisture intrusion. Experience indicates that the while specific grease 

formulations may have certain advantages over other formulations, the quality of the grease 

injection, specifically the removal of air pockets and voids, has been identified as the most 

critical factor for ensuring good corrosion protection with greased tendon systems [ORNL/TR 

6479]. It is important to also note that empirically, the US reactor fleet’s experience with greased 

tendon systems and corrosion protection has been very good as evidenced by the lack of cable 

degradation issues. Greased systems also offer the ability to remove a tendon, should a 

problem arise, where grouted systems almost certainly do not.

4.7 Tendon Duct Type and Effect on Corrosion Behavior

Research at the University of Texas showed that steel ducts perform much worse than do 

plastic ducts in terms of corrosion [Research Report 1405-4 1999, Research Report 1405-6 

2002]. The corrosion of the steel ducts was severe compared to that of the plastic ducts in the 

same conditions. Because of the corrosion of the steel ducts, large internal stresses were 

applied on the concrete specimens in which they were housed which led to concrete failure 

more often and more rapidly than when plastic ducts were used. The overall conclusion of the 

study was that plastic ducts performed much better than steel ducts with regard to corrosion.

More information on this topic has been found in reports in Europe (by the FIP and FIB Code 

Committees[CEB-FIP 1990, FIB-7 2000]. The following provides a summary of this information.

Provision of a corrosion resistant and leak tight encapsulation of the tendon can assure a very 

effective protection of the tendon. This concept has been used for many years for the protection 

of prestressed ground anchors. In the early 1990s, VSL introduced the corrugated plastic duct 

system, PT-PLUS, for grouted posttensioning (PT) tendons which together with suitable 

accessories such as connection details and anchorage caps provides a complete leak tight 

encapsulation of the post-tensioning tendons. The UK has made the encapsulation of tendons 

in plastic ducts compulsory in 1996. As a further step forward, the concept of verifying the leak 

tightness of the system was introduced. This verification is done by air pressure testing of the 

assembled duct and anchorage system. Pouring of concrete is only approved when the duct 

system is confirmed to be sufficiently air tight. Nevertheless, there is concern that plastic tendon 

ducts may not be appropriate for nuclear containment structures owing to the elevated 

temperature and radiation flux from the reactor. Additionally Freyssinet avoids the use of plastic 

ducts for nuclear containment structures because of the heat and pressure from the mass 

concrete curing in the thick concrete wall. 

If the encapsulation of tendons in plastic is supplemented with specific details at the 

anchorages, an “Electrically Isolated Tendon” (EIT) can be provided. In addition to the above 

mentioned advantages, an EIT allows monitoring of the provided encapsulation at any time 

during the design life of the tendon. 

A simple measurement of the electrical resistance between the tendon and the structure can be 

used to confirm the encapsulation of the tendon at any time. It can, in particular, be used to 

confirm the proper installation and the compliance of the tendon with the project specifications 

at the time of construction.

Encapsulation of tendons in plastic duct systems combined with EIT measurement was

introduced in Switzerland in 1993. Since that time more than 20 bridge structures have been 

built with this concept.
4-14

The positive experience with the concept has now led to the introduction of new guidelines for 

the protection of tendons in Switzerland. While still accepting some application of corrugated 

steel duct in a benign environment, these guidelines require encapsulation of tendons in plastic, 

in general. EIT is specified for a percentage of tendons to verify the encapsulation, and in 

general, for structures exposed to stray currents. It is not clear whether stray currents are an 

issue with NPP concrete structures, but the concept may have merit for nuclear structures. More 

research into the presence of stray currents in containment structures will help to evaluate the 

utility of EITs. 

4.8 Stress State as a Function of Corrosion

Various research cited in other sections describe the corrosion process that can occur in 

tendons, but there is a relatively little research describing how corrosion manifests itself within 

the tendon and surrounding concrete in terms of stress re-distribution and load-carrying. This 

chapter reports on a series of FEA studies which examine these questions, and make direct 

comparisons between how grouted versus ungrouted systems would behave.

For the study, we have focused on area deterioration, but for high pressure response of a 

PCCV, such assumptions could also simulate reductions in yield or ultimate stress of the 

tendons. In such cases, FEA simulations could be conducted by decreasing tendon area 

accordingly. In order to investigate the stress in the tendons and liner due to loss of tendon 

area, Model 1 of Standard Problem Exercise 3 [Akin, Sircar et al. 2013] has been reanalyzed 

with a range of corrosion assumptions, introduced while the tendons are under load. 

Corrosion is applied to one of the tendons at an anchor (at azimuth 270°) since anchor zones 

have been shown to have some susceptibility to corrosion. The opposite tendon is left 

uncorroded. Figure 4-2 shows the location of the corroded region. Figure 4-3 shows a 

representation of the loss of effective tendon area due to corrosion. This analysis is performed 

for both grouted and ungrouted tendons and the results are compared to see the effect of 

grouting on the redistribution of forces due to tendon corrosion. The friction between the tendon 

and sheathing is assumed to remain constant at 0.21.
4-15

Figure 4-2 Model 1 with Corroded Tendon at Anchorage

Figure 4-3 Reduction in Effective Area due to Corrosion
4-16

4.8.1 Loading

As in previous analyses, dead load and prestressing with jacking and anchorage is applied first. 

After anchorage, grouting is activated for the grouted case by locking the tendons to the 

concrete. For the ungrouted case, the tendons remain capable of sliding circumferentially with 

friction in the sheathing. Internal pressure is then increased to the desired pressure. Four 

separate pressure cases were analyzed: 1.0xPd (0.3725 MPa (54 psi)), 1.5xPd (0.55875 MPa 

(81 psi)), 2.5xPd (0.98125 MPa(142.32 psi)), and 3.0xPd (1.1175 MPa (162.08 psi)). The 

pressure was then held constant while the tendon was corroded at the anchorage by reducing 

the effective area of the tendon incrementally until only half of the tendon area remains. While 

such scenarios of corrosion occurrence while the structure is held at pressure are unlikely, the 

exercises were very illustrative as to force redistribution when corrosion occurs, and potential 

differences in response and performance between grouted and ungrouted systems.

Structural failure of the PCCV is controlled by the strain capacity of the tendons. For modeling 

purposes, the tendons were considered failed when they reached 3.8% strain. This criterion was 

evaluated offline by the analyst. 

4.8.1.1 Results

As the anchorage zone is corroded, the amount of force at the anchorage decreases, due to the 

reduction in tendon cross-sectional area. Similar to anchor set, the loss of force is largest at the 

anchorage and gradually the effects decrease due to friction. But for the ungrouted system, 

substantial azimuth is needed for the friction to ‘contain’ the perturbation in tendon strain. For 

the ungrouted case at 50% effective area, the loss in tendon force is observed almost half way 

around the vessel. Figure 4-4 shows the strain in the tendon profile as corrosion progresses for 

the ungrouted case. As the tendon with the corroded anchorage loses force, some load is 

transferred to the opposite tendon anchored at azimuth 0°. The increase in tendon strain in this 

opposite tendon is shown in Figure 4-5, starting from 270° to about 190° for 50% effective area. 

The deformation also causes the liner strains to increase, as shown in Figure 4-6. This increase 

occurs in the same region as the opposite tendon. The deformed shape plotted in Figure 4-7

shows the displacement of the vessel is largest at the location of the largest tendon force loss.
4-17

Figure 4-4 Strain in corroded tendon for ungrouted case @ 1.5xPd

Figure 4-5 Strain in uncorroded tendon (opposite) for ungrouted case @ 1.5xPd
4-18

Figure 4-6 Strain in liner for ungrouted case @ 1.5xPd

Figure 4-7 Displacement of liner for ungrouted case @ 1.5xPd

~160° affected
4-19

The results for the grouted condition are shown in Figure 4-8 through Figure 4-11. The transfer 

of force from the corroded tendon to the opposite tendon and liner are similar to the ungrouted 

case but the transfer occurs over a much shorter length due to the grouted condition, so the 

affected region is much more localized around the anchor zone. The loss of force in the 

corroded tendon and increase in the opposite tendon and liner occurs only for a few degrees of 

azimuth. Since the deformation is more focused at the anchorage, this also appears to cause 

more ovalization of the vessel.

For the grouted tendon case, the elevated liner strains extend across 65 degrees of azimuth 

once 50% of the tendon is corroded away. For the ungrouted case, at this corrosion, the 

elevated liners strains extend across only 20 degrees of azimuth. This demonstrates one of the 

primary differences in structure performance between grouted and ungrouted tendon PCCVs.

Figure 4-8 Strain in corroded tendon for grouted case @ 1.5xPd
4-20

Figure 4-9 Strain in uncorroded tendon for grouted case @ 1.5xPd

Figure 4-10 Strain in liner for grouted case @ 1.5xPd
4-21

Figure 4-11 Displacement of liner for grouted case @ 1.5xPd

The yield strain of the tendons and liner are 0.00816 and 0.00193, respectively. Since the 

corrosion of the tendon down to 50% of the effective area does not yield the tendon or liner at 

1.5xPd, larger internal pressures of 2.5xPd and 3.0xPd were applied to see the behavior of the 

model when liner and rebar yielding occurs. An internal pressure of 1.0xPd was also analyzed to 

see the behavior at the design pressure. 

Figure 4-12 through Figure 4-19 show results for 1.0xPd and indicate that the corrosion does not 

significantly affect the behavior of the vessel at the design pressure. 

Figure 4-20 to Figure 4-35 show results for the increased internal pressure for both grouted and 

ungrouted systems. For the larger internal pressure milestones, the tendon and liner strains 

follow the same trend as they did at 1.5xPd, but start at larger values.

It can also be noted that even though corrosion is only introduced at one azimuth for all of these 

case studies, the deformed shapes of the cylinder ring show that the entire circumference of the 

cylinder is affected. This is consistent with behavior observed for this cylindrical containment 

cross-section, dating all the way back to pre-test prediction analysis. The trend is explained by 

the cylindrical shape, and the frictional anchor effect which serves to resist the movement of the 

tendons (for the ungrouted case) associated with a local perturbation. Furthermore, the 

buttresses serve as a stiffness discontinuity in the circumference and create an effective plane 

of symmetry (from one buttress to the other) that results in a reflected response to the local 

perturbation. This behavior is in agreement with our understanding of the behavior of 

prestressed concrete structures, though specific experiments designed to assess the response 

of a PCCV to corrosion have not been performed. 

~65° affected
4-22

Figure 4-12 Strain in corroded tendon for ungrouted case @ 1.0xPd

Figure 4-13 Strain in uncorroded tendon for ungrouted case @ 1.0xPd
4-23

Figure 4-14 Strain in liner for ungrouted case @ 1.0xPd

Figure 4-15 Displacement of liner for ungrouted case @ 1.0xPd
4-24

Figure 4-16 Strain in uncorroded tendon for grouted case @ 1.0xPd

Figure 4-17 Strain in corroded tendon for grouted case @ 1.0xPd
4-25

Figure 4-18 Strain in liner for grouted case @ 1.0xPd

Figure 4-19 Displacement of liner for grouted case @ 1.0xPd
4-26

Figure 4-20 Strain in corroded tendon for ungrouted case @ 2.5xPd

Figure 4-21 Strain in uncorroded tendon for ungrouted case @ 2.5xPd
4-27

Figure 4-22 Strain in liner for ungrouted case @ 2.5xPd

Figure 4-23 Displacement of liner for ungrouted case @ 2.5xPd
4-28

Figure 4-24 Strain in corroded tendon for grouted case @ 2.5xPd

Figure 4-25 Strain in uncorroded tendon for grouted case @ 2.5xPd
4-29

Figure 4-26 Strain in liner for grouted case @ 2.5xPd

Figure 4-27 Displacement of liner for grouted case @ 2.5xPd
4-30

Figure 4-28 Strain in corroded tendon for ungrouted case @ 3.0xPd

Figure 4-29 Strain in liner for ungrouted case @ 3.0xPd
4-31

Figure 4-30 Strain in liner for ungrouted case @ 3.0xPd

Figure 4-31 Displacement of liner for ungrouted case @ 3.0xPd
4-32

Figure 4-32 Strain in corroded tendon for grouted case @ 3.0xPd

Figure 4-33 Strain in liner for ungrouted case @ 3.0xPd
4-33

Figure 4-34 Strain in liner for grouted case @ 3.0xPd

Figure 4-35 Displacement of liner for grouted case @ 3.0xPd
4-34

4.8.2 Conclusions

As the tendon at an anchor corrodes, the force in the tendon at the anchor decreases due to 

loss of effective area. This has the general effect of allowing increased radial displacement of 

the cylinder wall, and increased liner and rebar strains (increased hoop tendon strains as well). 

Grouting the tendons makes the force next to the corrosion decrease more, as shown in Figure 

4-36. As shown earlier, however, the extent of tendon force loss is very localized near the 

corrosion, whereas an ungrouted tendon has a larger force next to the corroded anchor but 

distributes the loss over a wider region. 

Figure 4-36 Strain in the corroded tendon adjacent to the corroded section which 

demonstrates the unloading of the tendon as the corroded section 

elongates

The maximum strain in the opposite tendon is plotted in Figure 4-37. This strain can occur 

anywhere along the circumference. At 1.5xPd, the maximum stress remains at the balance point 

due to anchor set and is not surpassed by the increase in stress at the corrosion, for both 

grouted and ungrouted cases. With 2.5xPd applied, the strain in the opposite tendon at the 

corrosion region becomes the maximum at an effective area of 70% in the corroded tendon for 

ungrouted condition. For the grouted condition, the opposite tendon strain at the corrosion 

region is the maximum as soon as corrosion occurs. For 3.0xPd with grouted tendons, the strain 

in the opposite tendon rises as soon as corrosion starts and increases rapidly. It is considered 

to be failed by 80% effective area.
4-35

Figure 4-37 Maximum Strain in Anywhere in Uncorroded Tendon which demonstrates 

the load transfer from the elongating corroding tendon to the uncorroded 

tendon

The liner strains are compared in Figure 4-38. For an internal pressure of 1.5xPd, grouting has 

little effect on the liner since the strains are so small. At 2.5xPd, the strains in the liner increase 

rapidly as corrosion progresses. Grouting the tendons yields the liner at 70% effective area of 

corroded tendon. For the ungrouted case, the liner yields at 60% effective area. By 50% 

effective area, the Maximum liner strains are the same for both grouted and ungrouted, but with 

more corrosion, the trend demonstrates that the ungrouted condition will have larger liner 

strains. When 3.0xPd is applied, grouted and ungrouted cases behave similarly until 80% 

effective area remains. After this, the grouted case liner strain increases drastically and fails 

before 70% effective area can be reached.

0.000

0.010

0.020

0.030

0.040

0.050

100 90 80 70 60 50 40

Strain

Effective Area Remaining (%)

Max Tendon Strain in Uncorroded Tendon

Ungrouted @ 1.0xPd

Grouted @ 1.0xPd

Ungrouted @ 1.5xPd

Grouted @ 1.5xPd

Ungrouted @ 2.5xPd

Grouted @ 2.5xPd

Ungrouted @ 3.0xPd

Grouted @ 3.0xPd
4-36

Figure 4-38 Maximum Liner Strain which demonstrates the degree to which the 

containment deforms in response to the corroding tendon

Figure 4-39 shows the radial displacement for the liner nodes as corrosion progresses. Similar 

to Figure 4-38, this plot serves to quantify the deformation of the containment cylinder in 

response to the loss of support from the corroding tendon.

Figure 4-39 Radial Displacement of Liner which shows the response of the containment 

to the loss of support from the corroding tendon
4-37

These conclusions are further quantified by the data in Table 4-2 and Table 4-3. This shows that 

maximum liner strains are somewhat higher for grouted systems, but the effects are more 

widespread for ungrouted systems.

Table 4-2 Maximum liner strain (%) at 2.5xPd

Corrosion Ungrouted Grouted

0% 0.2055 0.2294

10% 0.6750 1.2879

20% 1.4377 1.7384

30% 1.6291 1.8696

40% 1.8180 2.0400

50% 2.2089 2.2100

Table 4-3 Affected azimuth at 2.5xPd (degrees)

Corrosion Ungrouted Grouted

0% 0 0

10% 14 20

20% 19 20

30% 41 20

40% 44 20

50% 65 20

It should be noted here that the definition of “failure” for these studies is dictated only by tendon 

straining which corresponds to structural failure. Failure occurs when a tendon reaches its 

material limit of 3.8% strain [NUREG/CR-6810 2003]. In these studies, no consideration is given 

to liner strain driven liner tearing and leakage failure. For this series of studies, the only model 

case that “fails” is for 3.0 Pd with grouted tendons, and 80% effective remaining area.
4-38

4.9 Other Physical Conditions of PCCVs Affecting Corrosion and 

Durability

Previous studies by Sandia (Smith and Cherry [Cherry and Smith 1998, SAND2001-1762 2001]) 

have examined the causes and effects of steel component degradation in steel and concrete 

containments. These studies have included liner degradation, shell degradation (of steel 

containments), and tendon degradation of PCCVs. The work included FEA parameter studies 

and literature surveys. We summarize that work here in the context of the current study, and in 

the context of comparing potential impacts to grouted versus ungrouted tendons.

Regulations require that the prestressing tendons of PCCVs be inspected at 1, 3, and 5 years 

after the initial structural integrity test and every five years thereafter [Regulatory Guide 1.35 

1990, ASME 2010b, Regulatory Guide 1.90 2012]. Inservice inspections have, on occasion, 

found degradation of tendons. 

Degradation of prestressing can be associated with corrosion of tendons, corrosion of the 

tendon anchor or degradation of the concrete near the anchor, stress relaxation of the tendons, 

or from effects of high temperatures near the tendons. In addition to general corrosion, 

degradation of tendons can occur in the form of 

• pitting, 

• stress corrosion cracking, 

• hydrogen embrittlement.

Stress corrosion cracking occurs when there is the combination of high tensile stress, 

susceptible material, and an aggressively corrosive environment. 

Hydrogen embrittlement (also referred to as hydrogen assisted cracking) occurs when hydrogen 

is absorbed within the steel and interacts with micro-scale defects in the steel, creating a macroscale reduction in ductility. Hydrogen can be introduced into the steel during fabrication or as a 

byproduct of corrosion. 

A study documented by Shah and Hookham [Shah and Hookham 1998] provides detailed 

descriptions of each of these degradation types, and discuss tendon corrosion from hydrogen 

embrittlement, general corrosion, and stress corrosion cracking detected in a number of plants. 

They found that corrosion resulted from water accumulating near the ends/anchorages of the 

vertical tendons, and that was the result of poor construction practices. Tendons that have been 

stored on site for long periods with inadequate protection or that have not been properly 

protected once installed, but prior to the application of permanent corrosion protection, can 

experience corrosion. Shah and Hookham also reported that prestress losses detected at 20-

and 25-year tendon inspections exceeded those predicted for 40 years. So their reports express 

concern that loss of prestress is equivalent to loss of load-carrying capacity for the containment.

Ashar et al. [Ashar, Naus et al. 1994, Ashar, Tan et al. 1994] also discuss findings of 

containments with lower than expected prestressing levels. Plants identified were Calvert cliffs 

(some bearing plates were found to be depressed into the concrete), Bellefonte (failures 

occurred in the top anchor heads), Byron (four anchor heads of tendons failed between 1 and 

64 days after post-tensioning), Turkey Point (concrete cracking and grease leakage were noted 

at various locations on the dome – concrete delamination was found), and Crystal River (similar 

issues to Turkey Point). In one case cited, the levels of prestress in several hoop tendons, 

measured at three years after post-tensioning, were lower than levels predicted to occur after 40 

years. This is a severe detriment to aging performance. In another plant, vertical tendons were 
4-39

found with excessive losses during an inspection after about 13 years. The study suggested that 

contributing factors were: improper calibration of jacks during the initial post-tensioning, greater 

than expected losses from normal sources, and poor quality control.

An information notice released by the NRC discusses degradation of prestressing tendon 

systems in prestressed concrete containments at two plants. Between the two plants ten 

tendons were found to be degraded: one from broken wires, the other 9 from anchor-head 

failure. The lower than expected prestressing levels from higher than expected average 

temperatures around the tendons are also considered [NRC 1999].

Other researchers discuss the possibility that the minimum force in tendons may be lower than 

that calculated from the anchorage force, implying that the time-dependent losses along a 

tendon length could be higher than at the end anchorages. Steinberg gives numerical examples 

that shows variability of losses exceeds the losses expected by designers [Steinberg 1995]. 

Smith’s study [SAND2001-1762 2001] examined how much load-carrying capacity is lost and 

how the level of loss and pattern of degraded tendons affects the loss of capacity. It examined 

the effects of tendon degradation on the ultimate load-carrying capacity of the containment 

under a LOCA accident. A typical pressurized-water reactor (PWR) PCCV was analyzed using 

FEA with postulated tendon degradation. The containment model was similar to the Zion 

nuclear power plant containment. The level of tendon degradation and locations of tendon 

degradation were chosen independent of any known degradation at Zion NPP.

Current work postulating tendon degradation in a PCCV prototype represented by the Sandia 

1:4 Scale PCCV is described in detail in section 4.10. The current analyses extend the work 

performed by Smith by performing the FEA simulations with 3D global analysis (more readily 

completed with today’s computing power), and performing simulations and comparisons 

between ungrouted and grouted tendons. But Cherry and Smith’s work presents useful 

information about postulating the tendon corrosion. 

A literature review of corroded tendons was conducted to investigate whether the loss of 

ductility of degraded tendons could be quantified. Numerous references were found on 

degraded tendons. However, only a few offered quantitative data on the reduction of ductility. 

For corroded high-strength bridge wire, Barton et al. [Barton, Vermaas et al. 2000] found 

embrittlement and degradation of ultimate load that was in excess of what could be attributed to 

loss of section. The strain at ultimate load decreased by 40% from the uncorroded condition to 

the specimen with the maximum corrosion exposure time. They showed that hydrogen is 

definitely absorbed by corroding wire strands. Corrosion leads to increased embrittlement. They 

also measured that the embrittlement process cannot be reversed, thus showing that permanent 

microstructural damage is present.

In addition, Lopes and Simões [Lopes and Simões 1999] found that for prestresssing strands 

the number of bends for a reverse-bend test were reduced by as much as 50% for corroded 

specimens. Finally, Vehovar et al. [Vehovar, Kuhar et al. 1998] found that for corroded tendon 

specimens removed from a bridge structure, maximum elongation decreased from 6.0% down 

to 2.0%. In addition, the number of bends in a reverse-bend test decreased from 3.5 to fewer 

than 1. This loss of ductility was attributed to stress corrosion cracking.

Some disagreement in the literature has been found as to whether stress corrosion cracking 

and hydrogen embrittlement will significantly lower the failure stress. Cherry and Price [Cherry 

and Price 1980] suggested that hydrogen only assisted in the initiation of failure and that the 

failure would occur at nearly the ultimate tensile strength of the material. However, Parkins 
4-40

[Parkins, Elices et al. 1982] found that failures frequently occurred at stresses of 50% of the 

fracture stress. Parkins suggested that the reasons for the differences between their work and 

the work of Cherry and Price was that the test strain rates were not the same, the specimens 

were corroded in different solutions, and the Cherry and Price specimens were not notched.

Bergsma [Bergsma, Boon et al. 1977] documents results for prestressing tendons subjected to 

a corrosive environment that had a reduction in the reverse-bend values of 40 to 80% for the 

degraded specimens. Their results show that the effect of hydrogen embrittlement can depend 

on the solution to which the specimens have been exposed. The authors suggest that the 

corrosive environments are not necessarily realistic when compared to actual environments, so 

the results should perhaps not be used in service life predictions. But the results do show 

influence of hydrogen embrittlement on ductility of tendons and that failure levels can be 

considerably below the ultimate tensile strength of the tendons.

The research thus supports the conclusion that corrosion causes loss of ductility in the 

prestressing tendons. But it is not always clear whether the loss of ductility is due to general 

corrosion on the surface or due to hydrogen embrittlement or stress corrosion cracking. For 

cases of stress corrosion cracking and hydrogen embrittlement (such as can be found in high 

strength steels, such as prestressing tendons), the degraded material does not have the same 

properties as the undegraded material. The degraded material will have a decrease in ultimate 

load-carrying ability along with a loss of ductility (and not necessarily a loss of section). 

In Smith’s study [SAND2001-1762 2001] a 30° section of a typical PWR containment was 

modeled in Abaqus to study the effect of tendon degradation and prestressing loss on the 

response to internal pressurization. For some runs, the prestressing levels were lowered to 

simulate degraded tendons and in others, the cross-section was reduced to simulate corrosion. 

The strains in the tendons were examined but were not included in a failure criterion. The 

influence of premature tendon failure owing to lower than expected prestressing levels on the 

overall containment capacity was investigated using a liner failure criteria as the governing limit 

state for the containment. When tendon area was reduced or the tendon was removed (to 

simulate breakage) the pressure capacity was reduced, but when the prestressing force was 

reduced the containment pressure capacity was not significantly affected. 

4.10 FEA Modeling Tendon Corrosion for PCCVs and Predicted 

Effects on Performance

This section examines how corrosion of tendons can be modeled using FEA, and based on 

such modeling, reports how tendon corrosion could affect the global behavior of the PCCV. 

4.10.1 FEA Modeling

Three potential corrosion cases have been postulated. Using Model 3 from the Phase 1 work as 

a reference, certain areas of the vessel are assumed to suffer tendon corrosion, over time, and 

prior to occurrence of a severe accident. The authors suggest that a realistic way to simulate 

such field conditions is as follows:

1) Dead Load and Prestressing Loads are applied to the structure (when the structure is new);
4-41

2) Certain tendon groups/regions are down-sized to a smaller (corroded) tendon crosssectional area; this involves eliminating and replacing certain segments of tendon elements 

as an additional analysis step; structure equilibrium is allowed to be reached at the end of 

this step. Note that between Steps 1 and 2, the analyst could also introduce other “aging” 

conditions to the structure, such as creep of concrete, steel tendon relaxation, aging of 

concrete material properties (usually strength and stiffness increases), although these have 

not been introduced in the current study;

3) Severe accident pressure is applied up to structure system failure.

The research and findings of Cherry and Smith, discussed in subsection 4.9 add to the basis 

and relevance of introducing corrosion in this way. Smith’s research noted that certain forms of 

corrosion influence the tendons in additional ways, other than section loss; i.e., reduction in 

ductility, reduction in effective ultimate strength. But the authors view loss in section as the most 

likely scenario, and in fact, from the point view of effect on the global PCCV structure, loss in 

section is roughly equivalent to reduction in ultimate strength. A distinction of the current work 

compared to previous studies is the 3D global FEA analysis scope. Previous studies have used 

either sector models or axisymmetric models so that conclusions about behavior and failure 

mode were made on an axisymmetric basis. 3D global FEA (enabled by current computational 

facility) can show nonaxisymmetric behaviors and “early” failures associated with them. 

Capturing these nonaxisymmetric details and behaviors are critical for predicting the failureinitiating event or events. An axisymmetric model is incapable of identifying these local 

behaviors and can therefore only predict global failure. 

Three demonstration cases were performed. All cases consider grouted and ungrouted tendons. 

The first case applied corrosion to the outer vertical tendons along the bottom of the vessel. The 

second case applied corrosion to the hoop tendons adjacent to the equipment hatch. Both are 

locations potentially susceptible to such phenomena, and in fact Case 1 has been observed to 

have occurred in at least one U.S. plant. The third case applied corrosion to the horizontal 

tendons at an anchor zone, near Azimuth 270°, since anchor zones have been shown to have 

some susceptibility to corrosion. Only the tendons being jacked are corroded, the opposite 

tendons are left uncorroded. This analogous to the cases studied in subsection 4.8 with the 

simple two-tendon models.

Corrosion was applied as a reduction of tendon area by 60%, i.e., 6/10 of the tendon area is lost 

due to corrosion in the identified zones. For Corrosion Case 1 and 2, the affected areas are 

centered on the 0° azimuth, extend circumferentially 20°, and extend vertically, to form 

approximately a square zone of affected tendons. For Corrosion Case 3, the affected tendons 

are the tendon elements near 270° azimuth; the range extends vertically through the same 

elevations as Corrosion Case 2. Only one tendon element is corroded on each tendon for this 

case. A list of all of the Cases analyzed is provided in Table 4-4.

Figure 4-40 provides a general layout of where the corroded tendons were applied. The first 

case (Case 1) is the area hatched in green, which corresponds to the vertical tendons passing 

through the basemat extending up the wall. The second case (Case 2), in red, corresponds to 

the hoop tendons adjacent to the E/H opening. The third case (Case 3), in blue, corresponds to 

the hoop tendons at an anchor zone. Figure 4-41 through Figure 4-43 show the actual locations 

of the tendons corroded in the model. 
4-42

Figure 4-40 Postulated corrosion Zones

Corrosion 

Case 2

Corrosion 

Case 1

Corrosion 

Case 3
4-43

Table 4-4 List of Model 3 Analyses

Model 3

Corrosion First 

Then Pressure

Case 1 Ungrouted

Grouted

Case 2 Ungrouted

Grouted

Case 3 Ungrouted

Grouted

Case 1 & 2 Ungrouted

Grouted

Pressure First 

Then Corrosion

Case 2

1.5 x Pd Ungrouted

Grouted

2.5 x Pd Ungrouted

Grouted

Case 3

1.0 x Pd Ungrouted

Grouted

1.5 x Pd Ungrouted

Grouted

2.5 x Pd Ungrouted

Grouted

Figure 4-41 Corrosion Case 1 - Highlighted tendons indicate corroded areas along the 

bottom of the wall
4-44

Figure 4-42 Corrosion Case 2 - Highlighted tendons indicate corroded areas near the 

equipment hatch

Figure 4-43 Corrosion Case 3 - Highlighted tendons indicate corroded areas at anchor 

zone
4-45

4.10.2 Failure Criteria

Failure criteria for the PCCV are defined by strain limits for the tendons and liner, depending on 

the failure mode being studied. These failure strains are based on the stress-strain curve of the 

material. The stress-strain properties applied to the FEA models for tendons and liner are 

shown in Figure 4-44 and Figure 4-45.

Figure 4-44 Stress-Strain Properties for Tendon

Figure 4-45 Stress-Strain Properties for Liner

Based on the material curves, the strain limit that corresponds to tendon failure is ε = 0.038. 

Failure of the liner occurs when its ductility is exceeded. Locally, this ductility limit corresponds 
4-46

to strain greater than approximately 25%. However, based on the PCCV and other scale model 

test observations, and many liner detailed analyses, it has been shown that these conditions are 

reached at global liner strains on the order of 1%. The current work focuses primarily on tendon 

strain and tendon failure, but this 1% global strain limit should be kept in mind as a “leakage” 

milestone which will occur well before tendon failure strains are reached.

4.10.3 FEA Simulation Results

4.10.3.1 Corrosion Case 1 (Vertical Tendons near Wall-Base Juncture), Ungrouted

Figure 4-46 shows stresses in vertical tendons for Corrosion Case 1. After Anchorage no 

corrosion is present. After corrosion the stresses increase substantially in the corroded portion 

(due to the smaller cross-sectional area) and decrease elsewhere on those tendons, because 

the corroded area deforms significantly in response to the increased stress, thus reducing the 

deformation on the rest of the tendon. Similar patterns and conclusions can be drawn from the 

top views, Figure 4-47, before and after corrosion. Figure 4-47 shows that when the corrosion 

occurs on only one side (near one end of a hairpin tendon that is jacked from both ends), the 

effects of that corrosion dissipate over the dome, due to the frictional resistance between the 

tendon and the dome. The dome acts as a ‘friction anchor’ such that the corrosion on one side 

is not noticeable by the other end of the tendon.

A B

Figure 4-46 Stress in Vertical Tendons A) after Anchorage, and B) after Corrosion Case 1 

(Note Contours in units of psi)
4-47

A B

Figure 4-47 Stress in Vertical Tendons in Dome A) after Anchorage, and B) after 

Corrosion Case 1 (Note Contours in units of psi)

Figure 4-48 and Figure 4-49 show maximum principal stresses in the liner for Corrosion Case 1, 

after corrosion, and applied internal pressure equal to 1.5 and 3.3Pd, respectively. 

Figure 4-50 shows comparisons of radial displacement profiles cut through the Elevation 4.68m 

of the 1:4 Scale PCCV vessel (the elevation of the E/H). These show slightly larger radial 

displacements occurring near the 0-degree azimuth, and slightly lower radial displacements 

occurring at other azimuths. 

Figure 4-51 shows where liner strains (local zones near the E/H) have been extracted and 

compared in previous studies. Figure 4-52 show these liner strains for the current FEA. 

Comparing these to the strains at the same locations for the uncorroded case shows no 

significant differences between these strains.
4-48

Figure 4-48 Maximum Principal Strain in Liner at 1.5 x Pd, Corrosion Case 1, Ungrouted

Figure 4-49 Maximum Principal Strain in Liner at 3.3 x Pd, Corrosion Case 1, Ungrouted 

(Contours Modified)
4-49

Figure 4-50 Radial displacement comparisons between uncorroded, ungrouted and 

corrosion case 1 (Vertical Tendons at basemat-wall juncture, approximately 

0° azimuth), ungrouted

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0

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Ungrouted and Case 1 

Corrosion, Ungrouted Applied Prior to Pressure

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

Undeformed

PT

1 x Pd

1.6 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

Corroded 

Uncorroded

E/H

A/L

FW/MS

270° 90°
4-50

Figure 4-51 Locations for strains over selected gage length near E/H
4-51

Figure 4-52 Strains over selected gage length near E/H, corrosion case 1, ungrouted

4.10.3.2 Corrosion Case 1 (Vertical Tendons near Wall-Base Juncture), Grouted

The same analyses performed above for ungrouted tendons are re-analyzed with grouted 

tendons. To simulate grouting, the tendons are locked into the concrete (not allowed to slide) 

after the end of the prestressing analysis step (and after dead load). And with respect to 

corrosion, grouting is performed after the corrosion of the tendons, as it is assumed that 

grouting substantially prevents corrosion. This might not cover all possible scenarios of 

prototypical performance, but it does cover the scenarios of corrosion occurring due to poor 

QA/QC during on-site tendon storage and installation, and corrosive attack during the time 

period between tendon installation and grouting. Though rare, such situations have occurred in 

PCCVs (as described in subsection 4.9), and in fact, such situations are not so rare in other 

structure types, like bridges.

Since in the FEA, grouting is performed after corrosion, the stress and strain in the tendons at 

the beginning of the applied pressure step are the same for the grouted case as for the 

non-grouted case. The grouting of a corroded tendon may introduce other important physical 
4-52

phenomena such as reduced tendon bond strength that are not introduced here. These should 

be investigated further in future studies. 

The tendon stresses and strains are the same up through the beginning of pressure application 

(up through the grouting step), and they remain fairly similar during pressurization, so only the 

liner strains are provided.

The liner principal strains at initial and final pressure milestones are shown in Figure 4-53 and 

Figure 4-54. As before with ungrouted tendons, Case 1 does not exhibit significant increase in 

liner strains due to the tendon corrosion.

Radial displacements are compared in Figure 4-55. The differences are perceptible, but at most, 

are about 3%. At the azimuths nearest where the corrosion occurs, the corroded case shows 

the larger radial displacement.

Figure 4-56 shows liner strains near the E/H. Comparing these strains at the same location for 

the uncorroded case, locations 1 and 6 show slight decrease in strain for the corroded case. 

Location 9 shows slight increase in strain and the remaining locations shows no significant 

differences between these strains.

Figure 4-53 Maximum Principal Strain in Liner at 1.5 x Pd, Corrosion Case 1, Grouted
4-53

Figure 4-54 Maximum Principal Strain in Liner at 3.6 x Pd, Corrosion Case 1, Grouted

Figure 4-55 Radial displacement comparisons between uncorroded, grouted and 

corrosion case 1 (Vertical Tendons at basemat-wall juncture, approximately 

0° azimuth), grouted

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0

2000

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6000

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Grouted and Case 1 

Corrosion, Grouted Applied Prior to Pressure

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

3.4 x Pd

3.6 x Pd

Corroded

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

3.4 x Pd

3.6 x Pd

Uncorroded

E/H

A/L

FW/MS

270° 90°
4-54

Figure 4-56 Strains over Selected Gage Length Near E/H, Corrosion Case 1, Grouted

4.10.3.3 Corrosion Case 1 (Vertical Tendons near Wall-Base Juncture) Summary

Figure 4-57 and Figure 4-58 show whether either the liner or tendon has reached a failure 

condition, as mentioned in Section 4.10.2. 

Corrosion Case 1, Ungrouted, reaches failure at just above 3.0xPd. The liner strains at the 

pressure milestones shown do not show significant difference between this corroded case and 

the uncorroded case – the main difference is that generally speaking, the strains in the corroded 

case are slightly higher. The failure mode for the structure is failure of the corroded tendons. At 

first this observation seems to contradict the results from Section 4.8, but it is important to note 

that the hoop tendons modeled previously benefit from the angular friction as the tendon wraps 

around the circumference of the structure. This limits the load transfer to the corroded section. 

Conversely, for the ungrouted corroded vertical tendons, the deformation of the entire vertical 

tendon is concentrated in the corroded section, causing these tendons to fail at a lower internal 

pressure. 
4-55

Corrosion Case 1, Grouted, reaches 3.6xPd without any failure of tendons, which is similar to 

the uncorroded case. Liner strains at the pressure milestones, similar to the ungrouted case, 

does not show significant difference when compared to the uncorroded case. The main 

difference between the uncorroded and corroded cases is the increase in tendon strain. 

Figure 4-57 Maximum Liner Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted

0.000

0.002

0.004

0.006

0.008

0.010

0.012

0.014

0.016

0.018

0.020

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Strain

x Pd

Case 1 - Max Liner Strain

Uncorroded - Ungrouted

Uncorroded - Grouted

Corroded - Ungrouted

Corroded - Grouted
4-56

Figure 4-58 Maximum tendon strain comparisons between uncorroded and corroded, 

ungrouted and grouted

4.10.3.4 Corrosion Case 2 (Hoop Tendons near E/H), Ungrouted

Figure 4-59 through Figure 4-68 show stresses and strains in hoop tendons for Corrosion Case 

2. The figures show stresses for the tendons anchored at 90-degrees, separately from the 

tendons anchored at 270-degrees. After Anchorage (Figure 4-59, Figure 4-60), no corrosion is 

present, so the stresses are the same as for the ungrouted tendon analyses referenced in 

[NUREG/CR-6809 2003, NUREG/CR-6810 2003]. After corrosion (Figure 4-61 and Figure 

4-62), the stresses increase substantially in the corroded portion (due to the decreased 

sectional area) and decrease elsewhere on those tendons, because the corrosion reduces the 

overall forces in the affected tendons. 

Figure 4-63 through Figure 4-68 show tendon strains before and after corrosion, and at final 

pressure milestone, 2.8xPd. Prior to 2.8xPd, some hoop tendons begin failing (i.e., strains 

exceed 3.8%), and when this occurs, a progressive collapse mechanism ensues. As with the 

vertical tendon corrosion case and frictional effects in the dome, these figures again show that 

when hoop tendon corrosion occurs on only one side of the PCCV, the effects of that corrosion 

dissipate 180-degrees away, due to the frictional resistance between the tendon and the 

cylinder. The curvature of the cylinder acts as a ‘friction anchor’ such that the corrosion on one 

side is not noticable by the other end of the tendon, even though the tendons are ungrouted. 

This is an important result and observation in comparing/contrasting ungrouted tendon to 

grouted tendon behavior.

Figure 4-69 through Figure 4-71 show maximum principal strains in the liner for Corrosion Case 

2, at applied internal pressure equal to 1.5 and 2.8xPd, respectively. Figure 4-71 shows the 

same 2.8xPd result plotted on a deformed shape, with displacement magnification of 30. This 

0.000

0.010

0.020

0.030

0.040

0.050

0.060

0.070

0.080

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Strain

x Pd

Case 1 - Max Tendon Strain

Uncorroded - Ungrouted

Uncorroded - Grouted

Corroded - Ungrouted

Corroded - Grouted
4-57

shows that Corrosion Case 2 reaches apparent ‘failure’ at 2.8xPd. This can be compared with 

3.6xPd for the uncorroded case. 

Figure 4-72 shows comparisons of radial displacement profiles cut through the Elevation 4.68m 

(15.22 ft.) of the 1:4 Scale PCCV vessel (the elevation of the E/H). These show much larger 

radial displacements occurring near the 0-degree azimuth. Figure 4-73 shows liner strains at 

specific locations near the E/H. Comparing this to the strains at the same locations for the 

uncorroded case (see Ref. [NUREG/CR-6809 2003, NUREG/CR-6810 2003]) shows 

significantly larger strains for the corroded case.

Figure 4-59 Stress in hoop tendons anchored at 90° after anchorage, corrosion case 2 

(Note units of psi)
4-58

Figure 4-60 Stress in hoop tendons anchored at 270° after anchorage, corrosion case 2 

(Note units of psi)

Figure 4-61 Stress in hoop tendons anchored at 90° after anchorage and corrosion, 

corrosion case 2 (Note units of psi)
4-59

Figure 4-62 Stress in Hoop Tendons Anchored at 270° after Anchorage and Corrosion, 

Corrosion Case 2 (Note units of psi)

Figure 4-63 Strain in hoop tendons anchored at 90° after anchorage, corrosion case 2
4-60

Figure 4-64 Strain in hoop tendons anchored at 270° after anchorage, corrosion case 2

Figure 4-65 Strain in Hoop Tendons Anchored at 90° after Anchorage and Corrosion, 

Corrosion Case 2
4-61

Figure 4-66 Strain in hoop tendons anchored at 270° after anchorage and corrosion, 

corrosion case 2

Figure 4-67 Strain in hoop tendons anchored at 90° at 2.8 x Pd, corrosion case 2, 

ungrouted
4-62

Figure 4-68 Strain in hoop tendons anchored at 270° at 2.8 x Pd, corrosion case 2, 

ungrouted

Figure 4-69 Maximum principal strain in liner at 1.5 x Pd, corrosion case 2, ungrouted
4-63

Figure 4-70 Maximum principal strain in liner at 2.8 x Pd, corrosion case 2, ungrouted
4-64

Figure 4-71 Maximum principal strain in liner at 2.8 x Pd, showing deformed shape x30, 

corrosion case 2, ungrouted
4-65

Figure 4-72 Radial Displacement Comparisons between Uncorroded, Ungrouted and 

Corrosion Case 2 (Hoop tendons near the equipment hatch), Ungrouted

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0

2000

4000

6000

8000

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Ungrouted and Case 2 

Corrosion, Ungrouted Applied Prior to Pressure

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

2.8 x Pd

Corroded

Uncorroded

E/H

A/L

FW/MS

270° 90°
4-66

Figure 4-73 Liner Strains over Selected Gage Length Near E/H, Corrosion Case 2. 

Ungrouted

4.10.3.5 Corrosion Case 2 (Hoop Tendons near E/H), Grouted

As with Case 1 grouted, the grouting of the tendons for Case 2 occurs after corrosion has 

already taken place. 

Since in the FEA, grouting is performed after corrosion, the stress and strain in the tendons at 

the beginning of the applied pressure step are the same for the grouted case as for the 

ungrouted case. 

The final pressure application result for this case is shown in Figure 4-74 and Figure 4-75. The 

tendon stresses and strains are the same up through the beginning of the grouting step. Figure 

4-74 and Figure 4-75 show that as with the ungrouted tendons, large localized deformations 

occur due to the corroded tendons, and grouting makes the high strains more localized near the 

corrosion zones. Uncorroded portions of affected tendons are able to transfer higher loads to 

the vessel through the grout.

The liner principal strains at initial and final pressure milestones are shown in 
4-67

Figure 4-76 and Figure 4-77. Here the FEA does exhibit significant increase in liner strains due 

to the tendon corrosion, and the maximum local strains are larger for the grouted case than for 

the ungrouted.

Radial displacements are compared in Figure 4-78. The differences grow large near the 

corrosion zone, starting at pressure of 2.5xPd. Maximum radial displacements are similar for the 

grouted case as for the ungrouted case.

Figure 4-79 shows the local liner strains near the E/H. Comparing to Figure 4-73 for the 

ungrouted case, there is little difference between these local liner strains at a given internal 

pressure. However, there is a difference in the overall result of the simulation in that the grouted 

simulation ran to a greater internal pressure which resulted in larger overall liner strains.

Figure 4-74 Strain in hoop tendons anchored at 90° at 3.6 x Pd, corrosion case 2, 

grouted
4-68

Figure 4-75 Strain in hoop tendons anchored at 270° at 3.6 x Pd, corrosion case 2, 

grouted

Figure 4-76 Maximum principal strain in liner at 1.5 x Pd, corrosion case 2, grouted
4-69

Figure 4-77 Maximum principal strain in liner at 3.6 x Pd, corrosion case 2, grouted

Figure 4-78 Radial displacement comparisons between uncorroded, grouted and 

corrosion case 2 (Hoop tendons near the equipment hatch), grouted

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Grouted and Case 2 Corrosion, 

Grouted Applied Prior to Pressure

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

3.4 x Pd

3.6 x Pd

Corroded

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

3.4 x Pd

3.6 x Pd

Uncorroded

E/H

A/L

FW/MS

270° 90°
4-70

Figure 4-79 Local Liner Strains over selected gage length near E/H, corrosion case 2, 

grouted

4.10.3.6 Corrosion Case 2 (Hoop Tendons near E/H) Summary

Similar to Case 1, the four models of Case 2 are compared, uncorroded with ungrouted or 

grouted tendons and corroded with ungrouted or grouted tendons. Figure 4-80 through Figure 

4-82 summarizes the comparisons. 

Figure 4-80 compares the maximum between the four different combinations of grouted and 

ungrouted, corroded and uncorroded. For both corroded cases, the displacements are 

significantly higher than their corresponding uncorroded tendons. 

The maximum liner strains, shown in Figure 4-81, are slightly different between the corroded 

and uncorroded case – the strains in the corroded case are higher, especially in the vicinity of 

the corrosion for the ungrouted tendons. Figure 4-82 shows maximum tendon strain. The 

“structural” failure mode, irrespective of the liner, is failure of the corroded tendons. For 

ungrouted tendons, failure of tendons occurs at 2.1xPd. Grouted tendons fail around 3.4xPd. 

The “functional” failure mode will be liner tearing, and such will occur at correspondingly lower 

pressure. It should be noted that if certain aspects of corrosion (i.e., embrittlement) described in 

subsection 4.9 were to decrease the ductility of the tendons, the failure pressure would be 
4-71

further decreased. Further, the pressure difference between liner tearing and tendon failure 

would be narrowed. However, even if tendon ductility decreased from 4% to 2%, liner tearing 

would still occur first.

Figure 4-80 Maximum Radial Displacement Comparisons between Uncorroded and 

Corroded, Ungrouted and Grouted

-20.0

0.0

20.0

40.0

60.0

80.0

100.0

120.0

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Displacement (mm)

x Pd

Case 2 - Max Radial Displacement

Uncorroded - Ungrouted

Uncorroded - Grouted

Corroded - Ungrouted

Corroded - Grouted
4-72

Figure 4-81 Maximum Liner Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted

Figure 4-82 Maximum Tendon Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted

0.00

0.01

0.01

0.02

0.02

0.03

0.03

0.04

0.04

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Strain

x Pd

Case 2 - Max Liner Strain

Uncorroded - Ungrouted

Uncorroded - Grouted

Corroded - Ungrouted

Corroded - Grouted

0.00

0.02

0.04

0.06

0.08

0.10

0.12

0.14

0.16

0.18

0.20

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Strain

x Pd

Case 2 - Max Tendon Strain

Uncorroded - Ungrouted

Uncorroded - Grouted

Corroded - Ungrouted

Corroded - Grouted
4-73

4.10.3.7 Corrosion Case 1 and 2 Combined, Ungrouted

Figure 4-83 and Figure 4-84 show strains in hoop tendons for combined Corrosion Case 1 and 

2 after corrosion and at final pressure step for both tendons anchored at 90° and 270°. Similarly, 

maximum principal strains in the liner for Corrosion Case 1 and 2, after corrosion, and applied 

internal pressure equal to 1.0 and 2.3Pd are displayed in Figure 4-85 and Figure 4-86, 

respectively. The final pressure step is 2.3Pd, which is lower than both Corrosion Case 1 and 2, 

alone. 

Figure 4-87 shows comparisons of radial displacement profiles cut through the Elevation 4.68m 

of the 1:4 Scale PCCV vessel (the elevation of the E/H) between the uncorroded and corroded 

cases. These show much larger radial displacements occurring near the 0° azimuth. Figure 4-88

shows liner strains at specific locations near the E/H. Comparing the strains at the same 

locations for the uncorroded case shows larger strains for the corroded case.

Based on all plots, even though both Corrosion Case 1 and 2 play a role in a lowering failure 

pressure, results are similar to Corrosion Case 2. This indicates the hoop tendons near the E/H 

play a more significant role in lowering the failure pressure, compared to the vertical tendons 

near the wall-base juncture.

Figure 4-83 Strain in Hoop Tendon Anchored at 90° at 2.3xPd, Corrosion Case 1&2, 

Ungrouted (Final Step)
4-74

Figure 4-84 Strain in Hoop Tendon Anchored at 270° at 2.3xPd, Corrosion Case 1&2, 

Ungrouted(Final Step)

Figure 4-85 Maximum Principal Strain in Liner at 1.0xPd, Corrosion Case 1&2, 

Ungrouted
4-75

Figure 4-86 Maximum Principal Strain in Liner at 2.3xPd, Corrosion Case 1&2, 

Ungrouted (Final Step)
4-76

Figure 4-87 Radial displacement comparisons between uncorroded, ungrouted and 

corrosion case 1 and 2 (Vertical tendons at basemat-wall juncture and hoop 

tendons near the equipment hatch), ungrouted

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0

2000

4000

6000

8000

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Ungrouted and Case 1&2 

Corrosion, Ungrouted Applied Prior to Pressure

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

Uncorroded

Corroded

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.3 x Pd

E/H

A/L

FW/MS

270° 90°
4-77

Figure 4-88 Strains over Selected Gage Length Near E/H, Corrosion Case 1 and 2, 

UnGrouted

4.10.3.8 Corrosion Case 1 and 2 Combined, Grouted

Similar plots to the ungrouted case are shown below. Figure 4-89 through Figure 4-92 show 

strains in hoop tendons for Corrosion Case 1 and 2 after corrosion and at pressure milestones 

2.3 and 3.6Pd tendons anchored at 90° and 270°. Similarly, maximum principal strains in the 

liner for Corrosion Case 1 and 2, after corrosion, and applied internal pressure equal to 1.0, 2.3, 

and 3.6Pd are displayed in Figure 4-93 through Figure 4-95, respectively. 

Here the FEA reveals a decrease in tendon and liner strains due to the tendon corrosion, and 

the maximum local strains are smaller for the grouted case than for the ungrouted at 2.3Pd.

Radial displacements are compared in Figure 4-96. Similar to Case 2, the differences grow 

large near the corrosion zone, starting at pressure of 2.5Pd. Additionally, the local liner strains 

near the E/H are also similar to Case 2, shown in Figure 4-97.

-1.00E-03

2.00E-18

1.00E-03

2.00E-03

3.00E-03

4.00E-03

5.00E-03

6.00E-03

0 0.3925 0.785 1.1775 1.57

Strain over Gage Length

Pressure (MPa)

Strains Near Equipment Hatch

Loc. 1

Loc. 2

Loc. 3

Loc. 4

Loc. 5

Loc. 6

Loc. 7

Loc. 8

Loc. 9

Loc. 10
4-78

Figure 4-89 Strain in Hoop Tendons Anchored at 90° at 2.3xPd, Corrosion Case 1&2, 

Grouted

Figure 4-90 Strain in Hoop Tendons Anchored at 270° at 2.3xPd, Corrosion Case 1&2, 

Grouted
4-79

Figure 4-91 Strain in hoop tendons anchored at 90° at 3.6xPd, corrosion case 1&2, 

grouted

Figure 4-92 Strain in hoop tendons anchored at 270° at 3.6xPd, corrosion case 1&2, 

grouted
4-80

Figure 4-93 Maximum principal strain in liner at 1.0xPd, corrosion case 1&2, grouted

Figure 4-94 Maximum principal strain in liner at 2.3xPd, corrosion case 1&2, grouted
4-81

Figure 4-95 Maximum principal strain in liner at 3.6xPd, corrosion case 1&2, grouted

Figure 4-96 Radial displacement comparisons between uncorroded, grouted and 

corrosion case 1 and 2 (Vertical tendons at basemat-wall juncture and hoop 

tendons near the equipment hatch), grouted

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0

4000

8000

12000

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-16000 -12000 -8000 -4000 0 4000 8000 12000 16000

mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Grouted and Case 1&2 Corrosion, 

Grouted Applied Prior to Pressure

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

3.4 x Pd

3.6 x Pd

Corroded

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.6 x Pd

3 x Pd

3.3 x Pd

3.4 x Pd

3.6 x Pd

Uncorroded

E/H

A/L

FW/MS

270° 90°
4-82

Figure 4-97 Local Liner Strains over Selected Gage Length Near E/H, Corrosion Case 1 

and 2, Grouted

4.10.3.9 Corrosion Case 1 and 2 Summary

Figure 4-98 through Figure 4-100 show comparisons for Corrosion Case 1 and 2 and are 

summary plots of the results presented in subsections 4.10.3.7 and 4.10.3.8. 

Figure 4-98 compares the maximum radial displacement for the corroded tendon area near the 

equipment hatch and corresponds to Figure 4-87 and Figure 4-96. Figure 4-99 and Figure 4-100

compare maximum liner and tendon strains and summarize the numerous liner and tendon 

strain plots in subsections 4.10.3.7 and 4.10.3.8. The results follow similar trend as Case 2. 
4-83

Figure 4-98 Maximum radial displacement comparisons between uncorroded and 

corroded, ungrouted and grouted

Figure 4-99 Maximum liner strain comparisons between uncorroded and corroded, 

ungrouted and grouted

0.000

0.005

0.010

0.015

0.020

0.025

0.030

0.035

0.040

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Strain

x Pd

Case 1 and 2 - Max Liner Strain

Uncorroded - Ungrouted

Uncorroded - Grouted

Corroded - Ungrouted

Corroded - Grouted
4-84

Figure 4-100 Maximum tendon strain comparisons between uncorroded and corroded, 

ungrouted and grouted

4.10.3.10 Corrosion Case 3 (At Anchor Zone), Ungrouted

Figure 4-101 and Figure 4-102 show strains in hoop tendons for Corrosion Case 3 at final 

pressure step, 2.3Pd. As previously mentioned, the curvature of the cylinder acts as a ‘friction 

anchor’ such that the corrosion on one side is not noticeable by the other end of the tendon, 

even though the tendons are ungrouted. Furthermore, corrosion seems to only affect the 

tendons where the reduction in sectional area is applied, meaning corrosion does not greatly 

affect the nearby tendons anchored at the same buttress (Figure 4-102). This is most likely 

attributable to the significant stiffness of the buttress since the adjacent, non-corroded tendons 

are still well supported. Comparing Figure 4-101 and Figure 4-102, the 90° anchored tendons in 

Figure 4-101 are only very modestly affected by the corroded 270° tendons nearby (Figure 

4-102). The affected region in Figure 4-101 appears to be approximately 90° after which the 

friction anchoring effect overcomes the local discontinuity of the corroded tendons.

Figure 4-103 and Figure 4-104 show maximum principal strain in the liner for Corrosion Case 3, 

after corrosion, and applied internal pressure equal to 1.0 and 3.2Pd, respectively. This shows 

that Corrosion Case 3 is about to reach liner failure at 3.2Pd at the anchor zone where corrosion 

is applied.

Figure 4-105 shows comparisons of radial displacement profiles cut through the Elevation 

4.68m of the 1:4 Scale PCCV vessel (the elevation of the E/H). These show much larger radial 

displacements occurring near the anchor zone, at 270-degree azimuth. These affected tendons 

elongate in the region of highest tendon stress (near the anchorage) and large radial 

deformations accompany this tendon deformation. Comparing this to the strains at the same 

locations for the uncorroded case shows no significant differences between these strains since 

the corrosion seems to influence strains only in the immediate vicinity of corrosion.

0.000

0.010

0.020

0.030

0.040

0.050

0.060

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Strain

x Pd

Case 1 and 2 - Max Tendon Strain

Uncorroded - Ungrouted

Uncorroded - Grouted

Corroded - Ungrouted

Corroded - Grouted
4-85

Figure 4-101 Strain in Hoop Tendons Anchored at 90° at 3.2xPd, Corrosion Case 3, 

Ungrouted (Final Step)

Figure 4-102 Strain in Hoop Tendons Anchored at 270° at 3.2xPd, Corrosion Case 3, 

Ungrouted (Final Step)
4-86

Figure 4-103 Maximum Principal Strain in Liner at 1.0xPd, Corrosion Case 3, Ungrouted

Figure 4-104 Maximum principal strain in liner at 3.2xPd, corrosion case 3, ungrouted 

(final step)
4-87

Figure 4-105 Radial displacement comparisons between uncorroded, ungrouted and 

corrosion case 3 (Horizontal Tendons near anchorage, approximatel 260°

azimuth), ungrouted

4.10.3.11 Corrosion Case 3 (At Anchor Zone), Grouted

Similar to Case 1 and 2 grouted, the grouting of the tendons for Case 3 occurs after corrosion 

has already taken place. Since in the FEA, grouting is performed after corrosion, the stress and 

strain in the tendons at the beginning of the applied pressure step are the same for the grouted 

case as for the non-grouted case.

The tendon strains at internal pressures of 3.2 and 3.6Pd are shown in Figure 4-106 through 

Figure 4-109 for the tendons anchored at 90°and 270° azimuths. Unlike the ungrouted tendons, 

the deformation due to the corroded tendons is shared by both the 90°and 270° anchored 

tendons, allowing FEA of the structure to reach 3.6Pd. The FEA does exhibit increase in tendon 

strains due to the tendon corrosion, and in contrast to the previously presented uncorroded 

results, local maximum strains are smaller for the grouted case than for the ungrouted case.

The liner principal strains at pressure milestones 1.0 and 3.6Pd are shown in Figure 4-110 and 

Figure 4-111. 

Radial displacements are compared in Figure 4-112. The differences are similar to the 

ungrouted case, such that larger deformations occur near the corrosion zone. However, since 

the tendons are grouted, the radial displacements extend over a larger area, but with lower 

maximum deformations.

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Ungrouted and Case 3 

Corrosion, Ungrouted Applied Prior to Pressure

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

Undeformed

PT

1.1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.2 x Pd

Corroded 

Uncorroded

E/H

A/L

FW/MS

270° 90°
4-88

Figure 4-106 Strain in hoop tendons anchored at 90° at 3.2xPd, corrosion case 3, 

grouted

Figure 4-107 Strain in hoop tendons anchored at 270° at 3.2xPd, corrosion case 3, 

grouted
4-89

Figure 4-108 Strain in hoop tendons anchored at 90° at 3.6xPd, corrosion case 3, 

grouted

Figure 4-109 Strain in hoop tendons anchored at 270° at 3.6xPd, corrosion case 3, 

grouted
4-90

Figure 4-110 Maximum principal strain in liner at 1.0xPd, corrosion case 3, grouted

Figure 4-111 Maximum principal strain in liner at 3.6xPd, corrosion case 3, grouted
4-91

Figure 4-112 Radial displacement comparisons between uncorroded, grouted and 

corrosion case 3 (Horizontal Tendons near anchorage, approximately 260°

azimuth), grouted

4.10.3.12 Corrosion Case 3 (At Anchor Zone) Summary

Figure 4-113 shows the maximum radial displacement based on the profile cut through 

Elevation 4.68m (15.22 ft). Both ungrouted and grouted differ from the uncorroded cases, but 

the ungrouted tendons show more significant increases in displacement compared to both the 

uncorroded and corroded, grouted case.

Figure 4-114 and Figure 4-115 show if either the liner or tendon has reached a failure criterion. 

For Corrosion Case 3, Ungrouted, the system reached failure at just above 2.5Pd. The failure 

mode for the structure is failure of the corroded tendons. Past 2.5Pd, the strain in the tendons 

exponentially increase up to 11% at 3.2Pd where the analysis terminated. 

And for Corrosion Case 3, Grouted, the structure reached apparent ‘failure’ at 3.6Pd, which is 

similar to the uncorroded case. The liner strains at the pressure milestones shown are not 

significantly different between this corroded case and the uncorroded case – the main difference 

is that the strains in the corroded case are slightly higher. 

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Grouted and Case 3 Corrosion, 

Grouted Applied Prior to Pressure

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3 x Pd

3.3 x Pd

3.4 x Pd

3.6 x Pd

Corroded

Undeformed

PT

1 x Pd

1.5 x Pd

2 x Pd

2.5 x Pd

3.1 x Pd

3.3 x Pd

3.4 x Pd

3.6 x Pd

Uncorroded

E/H

A/L

FW/MS

270° 90°
4-92

Figure 4-113 Maximum radial displacement comparisons between uncorroded and 

corroded, ungrouted and grouted

Figure 4-114 Maximum liner strain comparisons between uncorroded and corroded, 

ungrouted and grouted

-0.005

0.000

0.005

0.010

0.015

0.020

0.025

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Strain

x Pd

Case 3 - Max Liner Strain

Uncorroded - Ungrouted

Uncorroded - Grouted

Corroded - Ungrouted

Corroded - Grouted
4-93

Figure 4-115 Maximum tendon strain comparisons between uncorroded and corroded, 

ungrouted and grouted

4.10.3.13 Summary of Corrosion Cases 1, 2, and 3

Figure 4-116 compares radial displacement at the E/H, from various analyses, to PCCV test 

Standard Output Location #14. 

0.000

0.020

0.040

0.060

0.080

0.100

0.120

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0

Strain

x Pd

Case 3 - Max Tendon Strain

Uncorroded - Ungrouted

Uncorroded - Grouted

Corroded - Ungrouted

Corroded - Grouted
4-94

Figure 4-116 Comparison of Radial Displ. at E/H – All Corrosion Cases, Grouted & 

Ungrouted vs. Uncorroded 

Table 4-5 through Table 4-9 tabulate the maximum tendon strain for each of the corrosion cases 

and grouting.

Table 4-5 Maximum Tendon Strain for Uncorroded Cases at Pressure Milestones

Uncorroded

Ungrouted Grouted

xPd Strain xPd Strain

0.0 0.70% 0.0 0.70%

1.0 0.71% 1.0 0.71%

1.5 0.71% 1.5 0.71%

2.0 0.70% 2.0 0.71%

2.5 0.77% 2.5 0.76%

3.0 0.96% 3.0 0.98%

3.3 1.21% 3.3 1.22%

3.4 1.37% 3.4 1.43%

3.6 1.74% 3.6 1.79%

-10.0

0.0

10.0

20.0

30.0

40.0

50.0

60.0

70.0

80.0

0.0000 0.3925 0.7850 1.1775 1.5700

Radial Displacement, mm

Pressure, MPa (grid divisions are multiples of Pd)

Standard Output Location #14. Azimuth: 334 Degrees, Elevation: 4.675 Meters,

Center of E/H

LST

Uncorroded

Corrosion Case 1 Ungrouted

Corrosion Case 1 Grouted

Corrosion Case 2 Ungrouted

Corrosion Case 2 Grouted

Corrosion Case 1&2 Ungrouted

Corrosion Case 1&2 Grouted

Corrosion Case 3 Ungrouted

Corrosion Case 3 Grouted
4-95

Table 4-6 Maximum Tendon Strain for Corrosion Case 1 at Pressure Milestones

Corrosion Case 1

Ungrouted Grouted

xPd Strain xPd Strain

0.0 2.41% 0.0 2.41%

1.0 2.47% 1.0 2.41%

1.5 2.51% 1.5 2.42%

2.0 2.52% 2.0 2.41%

2.5 2.94% 2.5 2.44%

3.0 3.68% 3.0 2.49%

3.3 7.56% 3.3 2.56%

3.4 2.54%

3.6 2.57%

Table 4-7 Maximum Tendon Strain for Corrosion Case 2 at Pressure Milestones

Corrosion Case 2

Ungrouted Grouted

xPd Strain xPd Strain

0.0 1.19% 0.0 1.19%

1.0 1.44% 1.0 1.21%

1.5 1.65% 1.5 1.23%

2.0 2.74% 2.0 1.33%

2.5 7.23% 2.5 1.71%

2.8 18.02% 3.0 2.52%

3.3 3.55%

3.4 3.97%

3.6 4.90%
4-96

Table 4-8 Maximum Tendon Strain for Corrosion Case 3 at Pressure Milestones

Corrosion Case 3

Ungrouted Grouted

xPd Strain xPd Strain

0.0 1.35% 0.0 1.35%

1.0 1.50% 1.0 1.36%

1.5 1.60% 1.5 1.40%

2.0 2.31% 2.0 1.51%

2.5 3.66% 2.5 1.71%

3.0 7.15% 3.0 2.18%

3.2 11.35% 3.3 2.64%

3.4 2.83%

3.6 3.57%

Table 4-9 Maximum Tendon Strain for Combined Corrosion Cases 1&2 at Pressure 

Milestones

Corrosion Case 1 & 2

Ungrouted Grouted

xPd Strain xPd Strain

0.0 2.41% 0.0 2.41%

1.0 2.47% 1.0 2.41%

1.5 2.50% 1.5 2.42%

2.0 2.74% 2.0 2.41%

2.3 4.10% 2.5 2.43%

3.0 2.56%

3.3 3.83%

3.4 3.94%

3.6 4.92%
4-97

General conclusions drawn from the FEA with tendon corrosion introduced to ungrouted and 

grouted tendons are:

• Reduction of tendon area provides a reasonable approach to analytically simulating 

corrosion, and solutions are obtainable;

• Expansion of the tendon due to corrosion was not considered nor was the associated 

cracking of grout and or concrete nor the loss of bond with the tendon;

• Reduction of area of about 60% is the apparent limit (based on this short study) for 

obtaining solutions in this way – area reductions more than 60% failed to converge when 

only subjected to the prestressing forces, i.e., the structure has difficulty redistributing the 

prestress when 60% is lost along a substantial bank of tendons;

• The vertical tendon corrosion case, ungrouted resulted in a reduction in ultimate pressure 

capacity (tendon failure capacity, NOT capacity defined by liner tearing) of about 10% 

(3.3Pd versus 3.6Pd);

• The hoop tendon near the E/H corrosion case, ungrouted resulted in reduction in ultimate 

pressure capacity of approximately 25%, with tendons beginning to rupture between 2.5 

and 2.8Pd, and pressure unsustainable beyond 2.8Pd. The reduction in capacity would be 

even greater if tendon embrittlement (reduction in ductility) were considered;

• Combining the vertical tendon and hoop tendon near the E/H corrosion case, ungrouted 

resulted similar to hoop tendon corrosion case, except with tendons failing between 2.0 

and 2.3Pd, and pressure unsustainable beyond 2.3Pd.

• The hoop tendons at the anchor zone corrosion case, ungrouted resulted in ultimate 

pressure capacity of approximately 10%, with tendons beginning to rupture at 2.5Pd, and 

pressure unsustainable beyond 3.2Pd;

• For all vertical and hoop tendon cases, the grouted tendon system resulted in ultimate 

pressure capacity nearly the same as for the corresponding uncorroded case. The main 

difference between the uncorroded and corroded case is the strains in the corroded cases 

are slightly higher.

• The most critical area of corrosion that causes early failure of the PCCV for both 

ungrouted and grouted cases occurs at the hoop tendons near the E/H. 

The overall conclusions of the 2012 FEA with corrosion studies are similar to those of Smith 

[SAND2001-1762 2001] (summarized in Section 4.9), but with some added observations about 

local behavior at the particular azimuth where corrosion is introduced. Reduced prestressing 

force (with associated degradation of prestressing tendons in PCCVs) showed that when the 

area of selected hoop tendons was reduced, there was a significant impact on the ultimate 

capacity of the PCCV. But when selected hoop prestressing tendons remained, but with loss of 

prestressing (Smith’s work), the predicted ultimate capacity was not significantly affected. This 

shows that the tendons’ presence as structural elements is more important to ultimate capacity 

than the prestressing effects on the concrete (although the response at lower pressures is 

affected). Concrete cracking occurs at lower levels for all cases. For cases where selected 

vertical tendons were analyzed with reduced prestressing or degradation of the tendons, there 

was not a significant impact on the ultimate load carrying capacity for the specific accident 

analyzed. For other loading scenarios (such as seismic loading) the loss of hoop prestressing 
4-98

with the tendons remaining could be more significant toward the ultimate capacity of the 

containment vessel than found for over pressurization.

4.10.4 FEA Simulation Results - Corrosion Applied After Pressure

In a manner similar to the approach used in subsection 4.8 for the two-tendon model, Corrosion 

Cases 2 and 3 were analyzed for Model 3, with the corrosion applied to the tendons after the 

internal pressure is applied. While this sequence of events is physically unrealistic, the 

approach is interesting and valid for several reasons. First, by reducing the tendon crosssectional area in steps, greater insight is gained into the level of corrosion that causes the 

structure to fail at a given internal pressure. Additionally, since the FEM simulations 

documented in this report are implicit calculations, the structure is in static equilibrium at the end 

of each step. This means that for each completed calculation step, the structural condition is 

valid for all the loads applied to it (internal pressure and reduction of prestressing force). 

As in previous analysis, dead load and prestressing with jacking and anchorage is applied first. 

Internal pressure is then increased to the desired pressure. Pressure is then held constant while 

the tendons corrode by reducing the effective area of the tendon incrementally until only half of 

the tendon area remains. For Corrosion Case 2, four analyses were performed; corrosion 

applied at 1.5 and 2.5Pd, ungrouted and grouted. For Corrosion Case 3, six analyses were 

performed; corrosion applied at 1.0, 1.5, and 2.5Pd, ungrouted and grouted. Plotted results are 

provided in the following sections.

4.10.4.1 Corrosion Case 2 Applied at 1.5 and 2.5xPd, Ungrouted and Grouted

Figure 4-117 through Figure 4-123 show the results of the corrosion case 2 simulations between 

ungrouted and grouted models with the corrosion applied at 1.5 and 2.5 x Pd.
4-99

Figure 4-117 Radial displacement comparisons between uncorroded, ungrouted and 

corrosion at 1.5Pd Case 2 (hoop tendons near the equipment hatch)

Figure 4-118 Radial displacement comparisons between uncorroded, ungrouted and 

corrosion at 2.5Pd Case 2

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Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Ungrouted and Case 2 Corrosion, 

Ungrouted Applied at 1.5xPd

Undeformed

PT

1 x Pd

1.5 x Pd

Undeformed

PT

1.0 Pressure

1.5 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

Corroded

Uncorroded

E/H

A/L

FW/MS

270° 90°

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Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Ungrouted and Case 2 Corrosion, 

Ungrouted Applied at 2.5xPd

Undeformed

PT

1 x Pd

2.5 x Pd

Undeformed

PT

1.0 Pressure

2.5 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

Corroded

Uncorroded

E/H

A/L

FW/MS

270° 90°
4-100

Figure 4-119 Radial displacement comparisons between uncorroded, grouted and 

corrosion at 1.5Pd Case 2

Figure 4-120 Radial displacement comparisons between uncorroded, grouted and 

corrosion at 2.5Pd Case 2

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Grouted and Case 2 Corrosion, 

Grouted Applied at 1.5xPd

Undeformed

PT

1 x Pd

1.5 x Pd

Uncorroded

Undeformed

PT

1.0 Pressure

1.5 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

Corroded

E/H

A/L

FW/MS

270° 90°

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Grouted and Case 2 

Corrosion, Grouted Applied at 2.5xPd

Undeformed

PT

1 x Pd

2.5 x Pd

Undeformed

PT

1.0 Pressure

2.5 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

E/H

A/L

FW/MS

270° 90° Corroded

Uncorroded
4-101

Figure 4-121 Maximum Radial Displacement Comparisons between Uncorroded and 

Corroded, Ungrouted and Grouted

Figure 4-122 Maximum Liner Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted
4-102

Figure 4-123 Maximum Tendon Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted

4.10.4.2 Corrosion Case 3 Applied at 1.0, 1.5, and 2.5xPd, Ungrouted and Grouted

Figure 4-124 through Figure 4-135 show the results of the corrosion case 3 simulations between 

ungrouted and grouted models with the corrosion applied at 1.0, 1.5, and 2.5 x Pd. Specifically, 

the corroded cases are represented by dotted lines in contrast to the solid lines for the 

uncorroded cases. Figure 4-124 through Figure 4-131 plot the radial displacement of the 

containment at the mid height while Figure 4-132 through Figure 4-135 present the results from 

multiple simulations as a function of corrosion.
4-103

Figure 4-124 Radial Displacement Comparisons between Uncorroded, Ungrouted and 

Corrosion at 1.0xPd Case 3

Figure 4-125 Radial Displacement Comparisons between Uncorroded, UnGrouted and 

Corrosion at 1.5xPd Case 3, Ungrouted

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Ungrouted and Case 3 Corrosion, 

Ungrouted Applied at 1.0xPd

Undeformed

PT

1 x Pd

Uncorroded

Corroded

Undeformed

PT

1.0 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

E/H

A/L

FW/MS

270° 90°

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-2000

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0

1000

2000

3000

4000

5000

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-6000 -5000 -4000 -3000 -2000 -1000 0 1000 2000 3000 4000 5000 6000

mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Ungrouted and Case 3 Corrorsion, 

Ungrouted Applied at 1.5xPd

Undeformed

PT

1 x Pd

1.5 x Pd

Uncorroded

Corroded

Undeformed

PT

1.0 Pressure

1.5 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

E/H

A/L

FW/MS

270° 90°
4-104

Figure 4-126 Radial Displacement Comparisons between Uncorroded, UnGrouted and 

Corrosion at 2.5xPd Case 3, Ungrouted

Figure 4-127 Radial Displacement Comparisons between Uncorroded, Grouted and 

Corrosion at 1.0xPd Case 3, Grouted

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Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Ungrouted and Case 3 Corrorsion, 

Ungrouted Applied at 2.5xPd

Undeformed

PT

1 x Pd

2.5 x Pd

Uncorroded

Undeformed

PT

1.0 Pressure

2.5 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

Corroded

E/H

A/L

FW/MS

270° 90°

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0

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2000

3000

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5000

6000

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mm

mm

Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Grouted and Case 3 Corrorsion, 

Grouted Applied at 1.0xPd

Undeformed

PT

1 x Pd

Undeformed

PT

1.0 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

E/H

A/L

FW/MS

270° 90° Corroded

Uncorroded
4-105

Figure 4-128 Radial Displacement Comparisons between Uncorroded, Grouted and 

Corrosion at 1.5xPd Case 3, Grouted

Figure 4-129 Radial Displacement Comparisons between Uncorroded, Grouted and 

Corrosion at 2.5xPd Case 3, Grouted

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mm

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Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Grouted and Case 3 Corrosion, 

Grouted Applied at 1.5xPd

Undeformed

PT

1 x Pd

1.5 x Pd

Uncorroded

Corroded

Undeformed

PT

1.0 Pressure

1.5 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

E/H

A/L

FW/MS

270° 90°

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mm

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Comparison of Deformation @ EL 4680 x 100 Between Uncorroded, Grouted and Case 3 Corrosion, Grouted 

Applied at 2.5xPd

Undeformed

PT

1 x Pd

2.5 x Pd

Corroded

Undeformed

PT

1.0 Pressure

2.5 Pressure

90% Eff Area

80% Eff Area

70% Eff Area

60% Eff Area

50% Eff Area

E/H

A/L

FW/MS

270° 90°

Uncorroded
4-106

Figure 4-130 Radial Displacement Comparisons between Uncorroded, Ungrouted and 

Corrosion at 2.5xPd Case 3, Ungrouted (Scale Magnified)
4-107

Figure 4-131 Radial Displacement Comparisons between Uncorroded, Grouted and 

Corrosion at 2.5xPd Case 3, Grouted (Scale Magnified)
4-108

Figure 4-132 Maximum Radial Displacement Comparisons between Uncorroded and 

Corroded, Ungrouted and Grouted

Figure 4-133 Average Radial Displacement Comparisons between Uncorroded and 

Corroded, Ungrouted and Grouted
4-109

Figure 4-134 Maximum Liner Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted

Figure 4-135 Maximum Tendon Strain Comparisons between Uncorroded and Corroded, 

Ungrouted and Grouted

0.000

0.001

0.002

0.003

0.004

0.005

0.006

100 90 80 70 60 50 40

Liner Strain

Effective Area Remaining (%)

Case 3 - Max Liner Strain

Ungrouted -

Corrosion @

1.0xPd

Grouted -

Corrosion @

1.0xPd

Ungrouted -

Corrosion @

1.5xPd

Grouted -

Corrosion @

1.5xPd

Ungrouted -

Corrosion @

2.5xPd

Grouted -

Corrosion @

2.5xPd
4-110

4.10.4.3 Summary of Corrosion Applied After Pressure Cases 2, and 3

Table 4-10 through Table 4-14 show the maximum tendon strain for each of the corrosion cases 

at different pressure milestones and grouting. The change in maximum strain between corrosion 

steps is generally negligible up to the point of tendon yield. After the tendon yields, the increase 

in maximum strain (at the corroded location) increases dramatically with additional area 

reduction.

Table 4-10 Maximum Tendon Strain for Corrosion Case 2 @ 1.5xPd

Corrosion case 2 @ 1.5xPd

Ungrouted Grouted

% Eff 

area

Strain % Eff 

area

Strain

100.0 0.71% 100.0 0.71%

90.0 0.71% 90.0 0.71%

80.0 0.71% 80.0 0.71%

70.0 0.71% 70.0 0.71%

60.0 0.72% 60.0 0.71%

50.0 0.86% 50.0 0.71%

Table 4-11 Maximum Tendon Strain for Corrosion Case 2 @ 2.5xPd

Corrosion case 2 @ 2.5xPd

Ungrouted Grouted

% Eff 

area

Strain % Eff 

area

Strain

100.0 0.78% 100.0 0.77%

90.0 0.78% 90.0 0.78%

80.0 0.80% 80.0 0.78%

70.0 1.03% 70.0 0.79%

60.0 2.00% 60.0 0.89%

50.0 3.09% 50.0 1.09%

Table 4-12 Maximum Tendon Strain for Corrosion Case 3 @ 1.0xPd

Corrosion case 3 @ 1.0xPd

Ungrouted Grouted

% Eff 

area

Strain % Eff 

area

Strain

100.0 0.71% 100.0 0.71%

90.0 0.71% 90.0 0.71%

80.0 0.71% 80.0 0.71%

70.0 0.71% 70.0 0.71%

60.0 0.76% 60.0 0.71%

50.0 0.93% 50.0 0.71%
4-111

Table 4-13 Maximum Tendon Strain for Corrosion Case 3 @ 1.5xPd

Corrosion case 3 @ 1.5xPd

Ungrouted Grouted

% Eff 

area

Strain % Eff 

area

Strain

100.0 0.71% 100.0 0.71%

90.0 0.71% 90.0 0.71%

80.0 0.71% 80.0 0.71%

70.0 0.71% 70.0 0.71%

60.0 0.77% 60.0 0.71%

50.0 0.97% 50.0 0.71%

Table 4-14 Maximum Tendon Strain for Corrosion Case 3 @ 2.5xPd

Corrosion case 3 @ 2.5xPd

Ungrouted Grouted

% Eff 

area

Strain % Eff 

area

Strain

100.0 0.77% 100.0 0.77%

90.0 0.79% 90.0 0.78%

80.0 0.92% 80.0 0.79%

70.0 1.16% 70.0 0.81%

60.0 1.67% 60.0 0.84%

50.0 2.47% 50.0 0.88%

For both Case 2 and Case 3, the vessel performs somewhat better with grouted tendons when 

the internal pressure is at 2.5xPd. At the smaller pressure of 1.0xPd and 1.5Pd however, it 

appears that ungrouted tendons have the advantage. For Case 2, Figures 10-89 to 10-92, and 

Case 3, Figures 10-101 to 10-104, it can be seen that at 1.5xPd, the grouted tendon models 

have larger displacements and liner strain. At 2.5xPd, the ungrouted tendons behave similarly to 

the grouted tendons until 60% effective area in the tendons remain; after that displacements 

and strains increase.

For Case 3, the displacements of this sequence of applying corrosion after internal pressure 

were compared to displacements when corroding before pressure. Figures 10-99 and 10-100 

show the displacements overlaid for comparison. When corrosion was applied first, 40% 

effective area was used. It can be seen that the deformed shape of corrosion first follows the 

trend of corrosion after pressure from 100% down to 50%. It can be concluded that once the 

desired pressure and corrosion is reached, the path taken to the final state does not matter 

significantly. 
4-112

4.11Transitioning From FEA to Probabilistic Study for Corrosion

Various efforts to formulate containment failure testing and analysis results into a probabilistic 

framework have been conducted over more than two decades in NRC and EPRI funded 

research, and by international organizations. Development, for example, of a probabilistic 

framework for predicting liner tearing and leakage was described in [EPRI NP6263-SD 1989, 

Dameron, Dunham et al. 1991]. And a final goal of the SPE-3 program (2011-2012) was to 

introduce a probability component to the leakage predictions versus pressure for PCCVs. 

Though more work is still needed to refine the approach, a similar framework as was laid out in 

the EPRI methodology summarized in [EPRI NP6263-SD 1989] was adopted in the SPE 3 

project. The method consists of the following steps:

1. Assume that prediction of liner strains from a global model, strain concentration factors, 

and liner ductility limits have a log-normal distribution.

2. Through statistical sampling of actual data, use of judgment, or an expert panel, assign 

parameters of i) Randomness, and ii) Uncertainty to liner strains from a global model, 

strain concentration factors, liner ductility limit, and leak rate formula versus leak area.

3. Apply the Randomness and Uncertainty parameters to each step of the liner tear 

prediction versus pressure, and sum these. This produces leak rate versus pressure with 

a lognormal distribution associated with every point on the leak rate versus pressure 

curve.

4. For any specific plant, other probabilistic aspects related to construction variations (liner 

thickness variations, weld quality, liner ductility variations, etc.) can also be introduced.

A similar philosophy could be applied to transitioning from FEA to probabilistic space for 

corrosion. The procedure would follow a development along the following lines.

Corrosion of tendons could be expressed by the loss in tendon steel ratios of the cylinder hoop 

tendons, the vertical (meridional) tendons, or the dome hoop tendons, respectively as:

, ,. ρρ ρ hc mc dc (4-2)

A refinement to ρhc could be pursued that takes into account the importance of the tendon’s 

proximity to the midheight of the containment, but for now, we will keep the approach simple in 

order to demonstrate a possible framework. First some approximations of containment 

performance should be made. The following presents such approximations for the Sandia 1:4 

PCCV Model structure (previously developed by Dameron in [NUREG/CR-6906 2006]).

The following definitions apply.

ρ = reinforcement ratio

ρhoop rebar = ρhr = Area of hoop reinforcement/gross concrete area

ρliner = Area of liner/gross concrete area

ρhoop tendons = ρht

tliner = thickness of liner = 0.16 cm (0.063 in.)

teq = equivalent concrete thickness or transformed section thickness (concrete section 

area with rebar and liner transformed by ratio of Young's Moduli)

t'eq = teq including rebar, liner, and tendons
4-113

tc = thickness of concrete wall = 32.5 cm (12.8 in.)

σo(concr) = compressive concrete stress after prestressing

R = Inside radius of cylinder = 538 cm (212 in.)

Erebar, Ec, Eliner, Es = Young's Moduli of rebar, concrete, liner, and general steel,

respectively 

εcr = Concrete cracking strain = 80x10-6

εry = rebar yield strain

σbarult = rebar ultimate strength

σtendonult = tendon ultimate strength (stress at 4.77% strain was used [NUREG/CR-6810 

2003])

4.11.1 Pressure at Which Cylinder Stress Overcomes Prestress, Po

Since there are three 16mm (0.63 in) hoop bars and five 13mm (0.512 in) bars every 45cm 

(17.7 in) (measured vertically)

ρhoop rebar x cm x cm

cmx cm = = 3 2 01 + 5 1 33

32 5 45

2 2

0 00865 . .

. . (4-3)

ρliner

t

cm

liner = = = 32 5

0 16

32 5 0 00492 . .. . (4-4)

( liner ) c

c

liner

hr c

rebar

eq t

E

E

E

E t = 1+ ρ + ρ (4-5)

tc

33

200(0.00492 (0.00865) 33

200 1 



 



 = + + (4-6)

t eq = 35.2cm (4-7)

There are four hoop tendons of area 3.39cm2 (0.53 in2

) in every 45cm (17.7 in.) wall segment.

ρhoop tendons x cm

x = = 4 3 39

32 5 45

2

0 00927 .

. . (4-8)

 0.00865 0.00492 0.00927 0.0228 ρtotal =++= (4-9)

In compression under tendon action,

0 ( ) 0.00927 950 tendon tendon σ ρσ concr =− × =− × MPa (4-10)

(Avg. prestress in hoop tendons including assumed losses)

σ o = − 8.83 1276 MPa psi ( ) (4-11)

Internal pressure to overcome prestress, Po is
4-114

( ) ( ) 0 8.83 35.3

0.580 84

538

eq

o

t

P MPa psi R

−σ − = = = (4-12)

4.11.2 Cylinder Hoop Cracking Pressure, Phc

Total equivalent t including tendons = t'eq

( ) 200 ' 1 1 0.0228 32.5 37.0

33

steel

eq total c

c

E

t t cm

E ρ

   =+ =+ =       (4-13)

Phc P t E

R o eq c cr x x x = + = +

′ − ε 37 0 33 000 80 10

538

6

0 580 . ( , ) . (4-14)

P MPa psi hc = 0.762 110.5 ( ) (4-15)

Including the pressure regime prior to hoop cracking, the hoop strain versus pressure response 

for the cylinder can be approximated by five piecewise-linear zones as shown in Figure 4-136. 

The boundaries of the zones are indicated by response events.

Figure 4-136 Approximate strain versus pressure response for cylinder hoop tendons 

in a PCCV

Pressure

Liner yield

Rebar yield

Tendon yield

Cylinder 

concrete hoop 

cracking

Cylinder Hoop Tendon Strain, εt

ZONE:

1

2

3

4

5

Ultimate 

Containment 

Failure
4-115

Continuing with hand calculation approximations,

4.11.3 Pressure at Liner Yield, Ply

Assuming the tendons have not yielded, the hoop stiffness, after the concrete cracks at an 

assumed value of 455 psi, is approximately that of elastic rebar, liner and tendons acting alone. 

Therefore,

( )

( ) 0.00228

200,000

Pr 455 = = = − = s

y

tota c s

y o

ry

t E E

P R

l

σ

ρ

ε

(4-16)

Solving, 

Pry = +

0 00228 0 0228 32 5 200 000

538 0 580 . ( . ) . ( , ) . (4-17)

P MPa psi ry = 1.21 175 ( ) (4-18)

Similarly, a straightforward approach can be constructed for calculating average hoop tendon 

strains (and by relationship, average hoop concrete, liner, and rebar strains) within each of 

Zones 1 through 5:

Figure 4-137 Tendon stress-strain curve

0

950 At P=O, 0.00475

200,000 

tho

th th

s

MPa

E MPa

σ

ε ε

  = = =     (4-19)
4-116

Additionally, at P=0, the average hoop concrete strain is

o

8.83 / 0.00027

33,000 c co c

MPa E

MPa

ε σ

  − = = = −     (4-20)

And the average hoop rebar and liner strains are compatible with the concrete, so:

00 0 0.00027 rl c εε ε = = = − (4-21)

Within any zone i, average hoop tendon strain equals: 

( ) 1 1 th thi i

eqi s

R P P

t E

ε ε =+ − − − (4-22)

The teqi values (below) are calculated by including the remaining contributors to hoop stiffness 

after each milestone. (Note that in this calculation, we are using a “transformed-section” 

approach, where we are transforming to equivalent steel thickness, thus the ratio of the moduli 

are applied to the concrete, not the steel.) For approximate purposes, after cracking, it should 

be assumed that concrete stiffness is zero, after yielding – rebar and liner stiffness are zero, 

and after tendon yielding, tendon stiffness is equal to Et yielded (non-zero) as shown in Figure 

4-137. ( ) 6 ~ 2 10 × psi or MPa 14,000

Calculating teqi

 1 cliner hoop rebar hoop tendons

s

E

teq tc E ρρ ρ

   

= ++ +        

(4-23)

teq tc 2 = ++ (ρρ ρ liner hoop rebar hoop tendons ) (4-24)

teq tc 3 = (ρ ρ hoop rebar hoop tendons + ) (4-25)

teq tc 4 = (ρhoop tendons ) (4-26)

5 4 t yield

s

E

teq teq E

−   =    

(4-27)

So th ε becomes a function of ρhoop tendons ( ) ρht and pressure. global liner ε − can be tied to th ε by the 

relation

 global lo th tho ε εεε =+− (4-28)

This is the same global ε used as the independent variable in [EPRI NP6263-SD 1989] in the 

formula
4-117

 p KBEglobal ε = (4-29)

So tendon corrosion could be introduced in much the same way as probabilistic liner strain is 

developed in [EPRI NP6263-SD 1989]. th ε becomes a variable that directly influences 

global liner ε − .

Corrosion could be introduced by:

1) Directly reducing ρht (by loss of section).

2) Indirectly reducing ρht by saying tendon yield strength is reduced (due to pitting or other 

means discussed in Chapter 9).

3) Reducing maximum strain (directly) if such damage as “embrittlement” is suspected.

We envision item 1 to be the most common parameter that would be encountered, so discuss 

this further here. The other parameters could be similarly introduced in a probabilistic 

framework. 

'

ht ( ) ht ρ ρ = corroded hooptendon ratio I C = − (4-30)

Where C is corrosion induced loss of Section Area as a fraction or per cent. Item three above 

might not influence global liner ε − directly, but would limit the maximum global liner ε − that could occur 

prior to a larger structural failure associated with tendon rupture. To create the probabilistic 

framework would then involve characterization of probability distribution and/or uncertainty 

bands for ρht and for C, and then convolving these into final leakage probability distributions.

Work in PRA for structures has been developed and refined over the past few decades, and 

many pioneering efforts in this field, particularly for nuclear power plant structures are 

referenced in [NUREG 1150 1990] as a basis for the PRA methods used in that document. The 

present development is a specific application of generally known PRA procedures with particular 

reliance on the work of Kennedy et al. [Kennedy, Cornell et al. 1980], Benjamin and Cornell 

[Cornell and Benjamín 1970], Ang and Tang [Ang and Tang 2006], and Clough and Penzien 

[Clough and Penzien 1993].

In the deterministic framework, the terms on the right-hand side of Equation (4-30) are 

considered to be best estimates or median values. Thus, in the probabilistic framework, Eq. (4-

30) is more appropriately recast as 

ˆ ˆ ' (1 ˆ) ht ht ρ ρ = − c (4-31)

and the probabilistic equation has terms

ˆ ht ht u r ρ ββ = p (4-32)

 ˆ u r c c = λ λ (4-33)
4-118

 ht ρ and C are now log-normally distributed random variables, and u γ , , R γ λu , λR , are lognormally distributed variables with unit median and logarithmic standard deviation βu and β R .

In order to construct Equations (4-32), (4-33) the right-hand terms must have reasonably 

lognormal dispersions and must be independent variables. The sources and inherent 

characteristics of each of these terms are discussed in detail in the development of the βu and 

β R .coefficients to follow. For now it suffices to say that the nature of the terms is similar to that 

of the various F factors used in seismic fragility curve development, i.e. factors characterizing 

calculated response and factors characterizing member strength. Therefore, the lognormal 

distribution assumption is deemed reasonable. 

A sample global ε curve showing graphical representation of the formula given in Equation (4-33)

is shown on Figure 4-138.

4.11.4 Summary of Randomness and Uncertainties

The assignment of randomness and uncertainty factors for the terms in the Corrosion formula 

are summarized below.

' 

0.15

0.05

, 

0.10

0.15

ht

u

R

u

R

corroded hoop tendon ratio

C Corrosion Factor

ρ

β

β

β

β

=

=

=

=

=

(4-34)

With the assumption of variable independency discussed and justified, the randomness and 

uncertainty factors are combined for the default values listed above as follows:

( ) 1 2222 2 0.15 0.05 0.10 0.15 0.15 βc = +++ = (4-35)

This gives the dispersion of the right-hand side of Eq. (4-30). 

It should be noted that some work is already available in the literature for assigning probabilities 

to prestress loss, for example in [Steinberg 1995], Steinberg describes and demonstrates a 

computer program developed to probabilistically evaluate prestress tendon forces over time.

The approximate method for calculating global hoop strains (both tendon strain and liner strain) 

is demonstrated by spreadsheet calculation and plotted in Figure 4-138.
4-119

Figure 4-138 Demonstration of approximate method for calculating global hoop tendon 

strain and liner strain

Theoretically, this liner strain can be assigned a probabilistic dispersion as illustrated in Figure 

4-139. (This is similar to the development in Reference [EPRI NP6263-SD 1989].)

Demonstration of application of hypothetical corrosion to hoop tendons, and the Influence on 

Global Liner Strain, is shown in Figure 4-140.
4-120

Figure 4-139 Sample global strain curve with confidence intervals

Figure 4-140 Demonstration of application of hypothetical corrosion to hoop tendons, 

and the influence on global liner strain

4.12 Detecting Corrosion in Grouted vs Ungrouted Systems

This chapter provides the result of the authors’ investigation related to detecting corrosion in 

grouted vs. ungrouted tendons. Particularly, questions are addressed as to what is the state of 

the art and where is the technology going.
4-121

Information about this topic is available from ACI’s 222R-01 report [ACI 1985]. (Specifically, 

Section 4.3.2). There is also a lengthy discussion on the state-of-the industry NDT methods 

available for monitoring prestressed tendons in the Task 2 Chapters – the companion report to 

this Task 3 report.

Identifying corrosion in structural concrete with prestressed tendons is generally more difficult 

than doing so with non prestressed reinforcement. Corrosion can occur in prestressed 

reinforcement without any outward signs of concrete damage and methods used to detect 

corrosion in reinforcing steel may not be applicable to prestressing steel. For example, it is not 

possible to detect corrosion of prestressing steel using corrosion potential or corrosion rate 

measurements since the steel is shielded by the sheath. The sheaths and/or ducts covering 

grouted prestressed tendons may be the most difficult of all structural steel systems to monitor 

for corrosion. Even when applying state of the art NDT methods, such as those discussed in 

another section of the report, it has been observed that non prestressed reinforcing steel may 

mask the behavior of prestressed reinforcing steel. This makes it difficult to interpret the data 

that impact echo or fiber optic technology, for example, may provide. 

Ungrouted prestressed tendons are generally much easier to asses for damage than are 

grouted systems. Although several methods are available, such as any of those discussed in 

the NDT section of this report, the best method for monitoring unbounded tendons is to remove 

the tendon and inspect it. Removing the tendon is currently the only method to completely 

assess the extent of corrosion along the tendon. ACI 423.4R provides guidelines for doing so. 

Again, it should be noted that NDT methods may run into the same issue as stated above for 

grouted tendons in that the reinforcing steel may mask the behavior of the unbounded tendon. 

In short, good NDT methods must seek to quantify the corrosion problem, not just identify it.

VSL also provides in depth information about inspection monitoring in [VSL Report Series 5 

2002]. The following is a summary of that information.

Post-tensioning tendons are structural elements essential for the safety, serviceability and 

durability of prestressed structures. Consequently, it is desirable to assess their behavior in 

existing structures. Such checks to detect possible defects or damages such as grout voids or 

tendon corrosion should be done by nondestructive or at least low destructive inspection 

methods and with minimum disturbance to structure operations.

4.12.1 Inspection and Monitoring Techniques

The inspection and monitoring methods listed below are focused on detecting existing grout 

voids, corrosion of the prestressing steel in progress, or even ruptured wires, strands or bars in 

tendons.

Georadar and Cover-meter: Experience with practical applications has shown that georadar is 

only suitable for the confirmation of the location of tendons. This is, however, often a 

prerequisite for a detailed tendon inspection. Whereas, under favorable conditions 

(no congestion of reinforcement) georadar allows the location of tendons to a depth of up to 500

mm, even a powerful cover-meter is generally not capable to detect ducts at concrete covers of 

more than 40 to 50 mm and again only if light reinforcement is present [Derobert and Coffec 

2001]. 

Potential Mapping: Whereas, potential mapping (measuring the potential field) is a powerful tool 

for finding corroded normal reinforcement, in case of tendons it is only successful under very 
4-122

favorable conditions (e.g. small concrete cover to the ducts and light normal reinforcement as 

they may exist in thin webs of precast beams).

Impact-Echo Method: Since 1983, the Impact-Echo Method has been under development 

primarily in the United States. It is stated that it can be used for detecting grout voids in tendons 

[Sansalone and Street 1995]. There is presently no standard method for assessing grout 

integrity for prestressing systems, and impact echo data must therefore be carefully scrutinized 

for its validity. Sansalone and Street recommend that impact echo data be calibrated at 

standard points to increase confidence in the results that are obtained. 

Applications in the United States have shown that under favorable conditions and in accessible 

areas, the Impact-Echo Method is able to identify grout voids. However, the method does not 

work with tendons in plastic ducts, which is not presently a widespread problem, but could be in 

the future if plastic ducts gain widespread acceptance for nuclear use.

Remanent Magnetism Method: The Remanent Magnetism Method was developed in Germany 

for detecting fractures in prestressing steel [Scheel and Hillemeier 1997]. The magnetizing and 

recording equipment has to be moved along the tendon path on auxiliary guidance rails and 

scaffolding fixed to the concrete surface.

Thus it allows localizing fractures in the accessible areas. The difficulty is disturbing magnetic 

signals originating from other embedded steel elements such as normal reinforcement, 

anchorage elements, duct couplers, steel plates, nails, etc.

Radiography: Today the application of radiography is limited to special cases. Even in France, 

where the method had formerly been widely used, it has practically disappeared. Apart from the 

high cost, another important reason is that most countries have national regulations for the 

protection of people, animals and the environment when applying radiography.

Whereas some of these regulations impose total evacuation (minimum distances depend on the 

intensity of the source; this generally means that all traffic has to be stopped in the area 

concerned), others ask for traffic suspension only when traffic cannot flow continuously.

Reflectometrical Impulse Measurement: Since about 1985, Time Domain Reflectometry was 

tried using known applications of coaxial telecommunication cables also for grouted tendons 

under the acronym RIMT. The method consists of sending high frequency impulses from an 

exposed anchorage through the tendon. By evaluating the recorded reflections it was hoped to 

detect anomalies along the tendon path. The results of research work done at the Institute of 

Technology in Zurich [Matt 2001] concluded that "the recorded signals do not contain 

information regarding the condition of the tendon but are artifacts of the measurement 

procedure. Thus RIMT cannot be used as a diagnostic technique for grouted tendons." In the 

opinion of the Swiss researchers, the high frequency of measurement artifacts confuse the 

interpretation of the signals so much that the technique can possess no practical value.

Ultrasonic Methods: Tests have shown that ultrasonic methods (transmission, reflection) for 

grouted tendons have very limited possibilities. Ultrasonic waves sent from a transmitter sitting 

on the end of the prestressing steel can detect anomalies only in special cases (e.g. only for 

smooth bars or wires) and only within a few meters from the tendon anchorage.

Acoustic Monitoring: To detect failures of prestressing steel by acoustic monitoring has been 

known for many years in fatigue testing of tendons and stay cables. Therefore, acoustic 

monitoring can be successfully applied in practice in equivalent situations such as for ungrouted

tendons and stay cables. Recently, trials have been carried out in Great Britain to assess 
4-123

whether the method can also be used for internal, grouted tendons. It is reported that these 

trials have been successful [Cullington, MacNeil et al. 2001]. It could be shown that a single wire 

fracture can be detected above the ambient noise level, distinguished from other acoustic 

events and even located in position. It is too early to say to what extent and in which situations 

acoustic monitoring will find its place in practical application. It can, however, be expected that 

the method will be restricted to special cases.

Other Methods: In the technical literature, further methods such as Thermography (infraredscanner) and tomography are also described.

In conclusion, a careful analysis of the suitability and limitations of these methods shows that 

none of them allow a full assessment of the conditions of a tendon. Some of them however, 

permit a partial assessment in favorable structural situations.

4.12.2 Engineering Approach to Tendon Inspections

While the above listed methods may allow a partial assessment of a structure and its tendons, 

the interpretation of the results is not easy and often ambiguous. However, there is one method 

which is quite basic and practical, and overall rates best in terms of information and 

interpretation.

This is the careful opening of tendon ducts by drilling into them, and subsequent visual 

inspection with an endoscope or similar device. Sometimes, it may be advantageous to open a 

window around the tendon location to obtain easier access for inspection and take samples for 

investigation.

Such careful opening permits the confirmation of presence of voids in the ducts at that particular 

location, and to investigate the grout (powder) collected during drilling for the presence of 

chlorides or other aggressive chemicals. These methods have been used successfully for many 

years for local isolated inspections. More recently, these methods have also been applied for 

the inspection of entire series of structures, see [Ganz and Vildaer] and have allowed a reliable 

assessment of these structures.

This method is particularly suitable if there is a reasonable doubt about the condition of a tendon 

at a particular location. Such doubt can be based on results of scanning methods presented 

earlier, or based on desk studies. Although this method is not non-destructive, the extent of 

intrusion is moderate, and is not considered harmful to the structure or tendon, if properly 

closed.

There are several publications to assist the engineer in such desk studies. In [Matt, Hunkeler et 

al. 2000] the authors conclude that the inspection engineer when assessing an existing 

structure should be aware of the possible hazard scenarios for post-tensioning tendons. Figure 

4-141 shows potential "weak points" in the case of a typical box girder bridge.
4-124

Figure 4-141 Typical inspection Strategy for a Box Girder Bridge

For each type of structure with its particular protection concept, the water, possibly chloride 

contaminated, can reach the prestressing steel in different ways. When assessing a 

posttensioned bridge, the study of the structural drawings, the construction and maintenance 

reports and the observations of the owner and his maintenance staff provide information 

regarding damaging actions and hazard scenarios. The key question is: Where does 

(aggressive) water get in contact with the structure and how does it flow off?
4-125

In addition, a thorough visual inspection (preferably after rainfall) of the concrete surfaces 

provides information on damage locations of the unstressed and stressed reinforcement and 

their location:

• Water flow, wet or moist areas

• Discoloration (e.g. rust stains)

• Spalling, delamination

• Cracks

• Honey-combing

• Concrete deterioration by freezing and thawing

• Joint leakage

The findings can then be substantiated by in-situ and laboratory investigations. Following these 

procedures in inspection and maintenance, potential corrosion damage of prestressing steel can 

be recognized and countermeasures taken.

4.12.3 New Developments in Monitoring

Post-tensioning systems have gradually evolved over the years. A significant step in the 

protection of tendons has been made with the introduction of the VSL PT-PLUSTM plastic duct 

system. This system was specifically developed for internal grouted post-tensioning, see [Ganz 

1997, FIB-7 2000]. In addition to the enhanced corrosion protection and service life of the 

tendon, lower and more reliable friction values, better fatigue performance, etc. this system can 

be fitted with anchorage details to provide a tendon which is electrically isolated from the 

surrounding structure. Thus, it is possible to check the integrity of the plastic duct encapsulation 

by measuring its electrical resistance against the surrounding concrete and normal 

reinforcement. Such testing also allows confirmation of the quality of tendon installation.

This new type of tendon, often called EIT (Electrically Isolated Tendon) has been applied since 

the early nineties primarily in Switzerland. Up to now, about 100 bridges have successfully been 

constructed using robust plastic ducts of which in over 20 bridges electrically isolated tendons 

have been installed. More applications are under execution or in the planning phase. The 

electrical resistance is periodically checked, and results are as expected.

It is important to note, that the protective envelope prevents the ingress of water and harmful 

substances. The grout, however, must still be of high quality.

Research at the University of Texas showed that steel ducts perform much worse than do 

plastic ducts in terms corrosion [Research Report 1405-4 1999, Research Report 1405-6 2002]. 

The corrosion of the steel ducts was severe compared to that of the plastic ducts in the same 

conditions. Because of the corrosion of the steel ducts, large internal stresses were applied on 

the concrete specimens in which they were housed which led to concrete failure more often and 

more rapidly than when plastic ducts were used. The overall conclusion of the study was that 

plastic ducts performed much better than steel ducts with regard to corrosion. More information 

on this topic has been found in reports in Europe (by the FIP and FIB Code Committees [CEBFIP 1990, FIB-7 2000]). 

Plastic ducts are also being used in U.S. Bridges, are becoming the material of choice by some

state DOTs, for example, Texas [FIB-7 2000, Matt, Hunkeler et al. 2000, Research Report 

1405-6 2002].
4-126

Potential corrosion advantages notwithstanding, plastic ducts are not used for nuclear 

applications owing to concerns related to radiation effects with the polymeric materials. Until 

such time as these concerns are addressed, metallic tendon ducts appear to be the only choice 

for prestressing systems in nuclear applications.

4.12.4 Repair of Tendons with Defective Grouting

As mentioned earlier, the careful opening of a tendon at questionable locations is currently the 

best method to verify its condition. Such probing allows determining possible defects and 

deterioration of a tendon, including its anchorages and couplers such as:

• Defective grouting (e.g. grout voids, grout segregation) and water access to the 

prestressing steel.

• Corrosion of the metallic duct, the prestressing steel, anchorages and couplers due to the 

ingress of water possibly contaminated by de-icing salts.

• Fretting corrosion of the prestressing steel due to fatigue.

• Corrosion of the prestressing steel due to stray currents [Bertolini, Carsana et al. 2007]

where the stray electrons change the electrochemical equilibrium and promote corrosion.

The inspection of a tendon by opening it locally is a low destructive method but has to be 

planned carefully. The planning should not only include the opening itself but also its closing 

after having carried out the inspection and the possibly required rehabilitation work of the 

tendon.

Based on non-destructive methods or desk studies, the engineer selects the tendon locations to 

be investigated. The exact tendon locations need then be indicated on the surface of the 

structure. This can be done based on post-tensioning shop drawings ("as built drawings") 

eventually supplemented with other methods, to confirm the location. It is recommended to 

involve the posttensioning specialist contractor to assist the engineer with system related 

questions. The closing of the tendon opening which has been created by either drilling or by 

cutting a window needs to be well prepared such that the tendon can be closed immediately 

after inspection and eventual repair - if possible on the same day.

The first step in the tendon inspection is to create access to the tendon duct or anchorage 

without damaging the duct or prestressing steel. The access needs to be kept as small as 

possible, at least initially. The following methods have been successfully used:

• Drilling of a core of 50 to 80 mm (1.97 to 3.15 in.) diameter. The drilling machine can be 

equipped with an automatic switch-off when the core touches the metallic duct.

• Removal of the concrete cover with an electric pick hammer. The concrete just adjacent to 

the tendon duct should be removed preferably by hand with light equipment. The tendon 

duct can then be opened for tendon inspection in the following steps:

• Cutting of the duct by hand with small, hand-held equipment such as disc cutter and flat 

chisel, and removal of the cut duct section. The duct opening is preferably kept smaller 

than the access in the concrete.

• Small samples of grout can then be removed for analysis of the chloride content. 

Typically, a few grams of grout per location of sampling are sufficient.
4-127

• If the duct is partially or completely without grout, visual inspection is possible and photos 

can be taken with an endoscope.

• If the prestressing steel is corroded, samples of corrosion products can be collected for 

analysis in a qualified laboratory to determine the type of corrosion. 

Above methods to gain access to the tendon are illustrated in Figure 4-142. 
4-128

Figure 4-142 Gaining Access to a Tendon [VSL Report Series 5 2002]

Core drilling (with automatic switch-off)

Reinforcement

Automatic switch-off

Tendon Core drilling 

machine

Chiseling by hand

Chiseling with 

electric pick hammer

Reinforcement

Cutting of duct with 

chisel or disc cutter

Reinforcement

Fold back 

and cut

Collect grout for 

analysis

Duct with defective 

grouting

Camera

Endoscope with rigid

or flexible fiber probe Reinforcement

Chiseling

Opening of the duct Investigation of duct with endoscope
4-129

4.12.5 Grouting of Voids

Before starting the grouting of existing voids, the exposed prestressing steel must be carefully 

cleaned by to prevent pitting corrosion or chloride contamination inside the duct. In order to 

select the appropriate grouting procedure, it is necessary to determine the geometrical 

characteristics of the detected void (length, cross-section, volume etc.). In case of a larger void, 

vacuum assisted grouting is recommended. In special cases, the vacuum technique can also be 

used to measure the volume of the void and thus, determine its extent along the tendon.

The vacuum pump reduces the air pressure inside the duct to a certain subatmospheric 

pressure (e.g. about 80% of the atmospheric pressure). The procedure is then automatically 

reversed and the air flowing back into the duct is measured and recorded.

In order to determine the precision of the applied equipment, preliminary tests are 

recommended for calibration.

In principle, only cementitious, alkaline materials should be used for void filling.

In case of very small voids, these can be patched by using a suitable mortar (thixotropic, if 

required). Tremie grouting can be applied with voids that are still comparatively small (maybe 

over a length of about one meter).

For larger voids several meters long, vacuum grouting is recommended using the same material 

as for new grouting. At the end of the grouting operation, the pressure should be increased 

typically 1-3 bars and held constant for about 1 minute.

The effectiveness of the chosen method should be tested beforehand. Figure 4-143 shows 

vacuum injection equipment which permits measurement of void and grout volume. The 

advantage of the vacuum method is that only one access to the void at any location is required. 

In general, this can be the borehole which has been made for the inspection of the tendon and 

for taking samples to determine the chloride contamination.

A comparison of the previously measured void volume and the injected grout confirms the 

success of the procedure. In case of discrepancies, it may be necessary to make checks by 

additional boreholes.
4-130

Figure 4-143 Vacuum Control and Grouting Equipment for Tendons with Defective 

Grouting

4.12.6 Closing of the Tendon

In the following, four possibilities are given for the repair of tendon openings depending on 

access, see Figure 4-144. In most cases, due to the presence of normal reinforcement, it is not 

possible to provide an additional protection by installing a half duct. Where the conditions are 

favorable, replacement of the removed duct section should, however, be considered. The 

placing of repair concrete or mortar on to the tendon grout has to generally be made "wet-towet" to assure optimum bond. 
4-131

Figure 4-144 Closing of Tendon Openings [VSL Report Series 5 2002]
4-132

1) Accessing the tendon from above:

• Roughening and cleaning of concrete surface

• Wetting of concrete surface

• Filling of duct and covering the vicinity of the duct with a minimum of 40 mm of 

cementitious grout

• Filling of the remaining space of the opening with a shrinkage compensated 

cementitious repair mortar in several layers in accordance with the instructions of the 

mortar supplier.

2) If on an overhead surface:

• Concrete with a shrinkage compensated cementitious repair filler in several layers in 

accordance with the instructions of the filler supplier. In case of large voids inside the 

duct, the duct can subsequently be vacuum injected through a hose placed into the filler.

3) If on a vertical surface:

• Placing and sealing of formwork over the opening

• Partial filling of the tendon opening and duct with a cementitious grout

• Removal of the formwork

• Filling of the remaining opening with a shrinkage compensated cementitious repair filler 

as in Item (2) above. 

• Alternatively, a thicker cover can be formed and filled with cementitious grout. A 

minimum of 40 mm on a roughened concrete surface is recommended.

4) Pouring the concrete back:

Alternatively to the above methods (1) to (3), the tendon opening can be poured back with 

concrete. This is particularly suitable, if the openings are large, e.g. in bridge box girders:

• Roughening, cleaning, and wetting of concrete surface 

• Placing and sealing of formwork

• Pouring back the opening with repair concrete. In all the above cases (1) to (4) it may be 

considered to provide an eventual protection of the concrete surface against ingress of 

humidity or chlorides with special surface protection systems.
4-133

4.13 Summary of Grouted and Ungrouted Tendons

In [VSL Report Series 5 2002], VSL has provided their summary of pros and cons of grouted 

versus ungrouted systems, and they comment on some differences in tendon system 

performance (i.e., grouted versus ungrouted tendons).

“In our opinion, there is no one type of tendon which answers all requirements, and it is 

up to the engineer to select the type of tendon best suited to a particular project and 

construction method.”

Advantages of grouted tendons (as recommended by VSL [VSL Report Series 5 2002]):

• Provision of active corrosion protection: The prestressing steel is actively protected, i.e. 

passivated, against corrosion through the alkaline environment provided by the 

cementitious grout. To initiate corrosion prevention, steel first needs to be depassivated.

• Provision of bond of the tendon to the structure: Bond allows an increase of the 

prestressing force in a cracked section after decompression, and permits the prestressing steel to reach the yield or even ultimate strength. This has effects on the 

strength of a section, on the crack distribution in the prestressed member, and on the 

energy dissipation of the member [VSL 1992]. Bond also has beneficial effect on the 

redundancy of a prestressed member. A local defect in the tendon remains local, i.e. the 

tendon force is not affected over the entire tendon length.

• Cost effectiveness: Cementitious grout is a cost effective injection material for which 

long and good experience exists. The compatibility of cementitious grouts with 

prestressing steel is well proven over a long period of time. 

Advantages of ungrouted tendons (as recommended by VSL [VSL Report Series 5 2002]):

• Future adjustment of prestressing force: Prestressing forces of ungrouted tendons can 

theoretically be adjusted at any time during the design life of a structure. However, all 

necessary tendon details for later stressing need to initially be provided such as access 

and clearance for jacks, and sufficient extra length of prestressing steel to connect the 

jack to the strand. While re-stressing of tendons is a justifiable concern when long term 

losses due to creep and shrinkage of concrete, and relaxation of prestressing steels, 

were not yet well understood, this is no longer the case today. The authors are not 

aware of any recent case where re-stressing of a tendon was necessary due to 

excessive losses of tendon force. We would like to give a word of caution because retensioning of a tendon, initially stressed to 70-80% of its strength, at some time during 

the design life of the structure is certainly not an easy task. Hence, if an increase in 

prestressing force is ever required, the best option seems to be to provide additional 

tendons to the structure. A number of recent standards such as AASHTO, [AASHTO 

2003], actually require new structures to be detailed for the addition of future external 

tendons to potentially increase the prestressing force to accommodate potential increase 

of loads or excess loss of tendon force. According to these standards, anchorages and 

deviation details need to be provided to allow addition of a fixed number of tendons, e.g. 

2 per section, or of a given percentage (AASHTO: 10%) of the initial prestressing force. 

This procedure keeps the initial investment to a minimum, and greatly facilitates the 

future addition of tendons to the structure, if ever needed.
4-134

• Facilitated inspection of tendon: Since ungrouted tendons are placed externally to the 

structure, access to the tendon for inspection is facilitated over a substantial portion of 

the tendon length. Such access is not usually available near the anchorages and/or at 

tendon deviation points where such tendons often are anchored or deviated in massive

diaphragms. While access to the tendon is facilitated, inspection of the prestressing steel 

inside the tendon or bundle of prestressing strands is not necessarily provided. Hence, 

special inspection or monitoring devices still need to be used to collect information on 

the actual performance and durability of the steel.

• Replaceability of tendons: Ungrouted external tendons may be replaced at any time 

during the design life of a structure. Replacement is preceded by either de-tensioning of 

the tendon if the necessary tendon details have been initially provided, or by gradual 

cutting of the tendon according to specific procedures adapted to the particular site and 

tendon type. The actual removal of the tendon is then possible if appropriate details 

have been provided initially at anchorages and deviation points. Installation of a new

tendon can then follow. The authors are of the opinion that tendon replacement should 

only be considered if there is a significant risk of unexpected tendon failure with 

consequential damage or risk to persons. In all other cases and in particular if the 

structure can accept additional prestress, rather addition of new than replacement of 

existing tendons should be considered.

Such favorable conditions to avoid replacement exist, in particular, for grouted tendons 

in structures with sufficient concrete dimensions.

In the late 1970s, when some of the currently operating U.S. nuclear fleet was still being 

designed, Oak Ridge National Laboratory performed studies of post-tensioning systems for 

PCCVs, and published a similar list of pros and cons, but the list covers a few additional items 

particularly germane to NPPs [ORNL/TR 6478 and ORNL/TR 6479].

Below are pros and cons of grouted/ungrouted tendons per the Oak Ridge studies [ORNL/TR 

6478]:

Present practice in the United States for fabrication of PCCVs is to use post-tensioned 

prestressing tendons that are ungrouted. This decision has been based largely on one or more 

of the following arguments in favor of ungrouted tendons:

1) Tendon loads may be periodically monitored with retensioning, as required.

2) Tendons may be removed, inspected for corrosion, and replaced. If necessary.

3) Poor grouting practices can lead to an acceleration of the corrosion process.

4) Tendon stresses are distributed along the full length of the tendon, which can lead to 

more ductile behavior than with grouted systems.

5) Corrosion-inhibiting compound reduces friction losses because it acts as a lubricant.

Proponents of grouted tendon systems feel that grouted tendons provide superior performance 

at reduced cost. Arguments cited for using grouted tendon systems are as follows:

1) Performance is superior for flexure members with ultimate load increases of up to 50% 

and cracking load increases of up to 10 to 15% relative to ungrouted tendon companion 

specimens.
4-135

2) Effective grouting has been shown to provide an easy technique for corrosion protection 

with possible avoidance of periodic monitoring and maintenance of the corrosioninhibiting medium.

3) Crack control is improved. More cracks form, but average crack widths are smaller so 

that strains transferred to the liner at crack locations are significantly reduced.

4) Dynamic effects are eliminated or significantly reduced if a tendon were to fail, with 

prestressing force lost only in the vicinity of the failure; that is, if an anchorage or tendon 

fails, effects are localized and overall strength insignificantly affected.

5) Grouting provides conservatism in seating and overall anchorage, particularly under 

fluctuating load conditions such as occur with an earthquake; that is, reduction in 

anchorage efficiency may result without a reduction in ultimate load.

On the issue of cost comparison, it must be noted that corrosion protection of ungrouted 

tendons in PCCVs during their anticipated 30- to 40-year service life is generally provided by 

encapsulating the tendons in organic-petrolatum-based greases and waxes containing corrosion 

inhibitors. Inservice inspections (visual, prestress monitoring, and material tests) should be 

performed one, three, and five years are required after the initial containment structural integrity 

test and every five years thereafter. During the operating life of the vessel, this may amount to 

as many as ten inspections. These inspections are performed under a surveillance contract and 

the costs can be substantial. Presently inspection and monitoring for grouted prestressing 

systems is in a state of development in terms of specific techniques and frequency. As such, 

estimating the specific costs for monitoring grouted systems is quite difficult, but there is no 

reason to think the cost would be significantly different for these systems.

5-1

5 SUMMARY AND CONCLUSIONS

Throughout the world, the regulation of grouted tendons for use in nuclear power plants is still 

being developed. Much of the world has referenced the US NRC’s regulatory guidance, 

standards, and documents for use in their own countries. Other countries such as Belgium, 

Canada, and France have more operating plants with grouted tendon systems and have 

developed their own regulatory practices. This document serves to consolidate this (and other) 

experience. Some plants with grouted tendons have seen many years of service without 

reported corrosion damage of note to date in the grouted tendons, however corrosion is difficult 

to measure directly for grouted tendon systems. 

Despite the lack of reported corrosion damage in PCCVs with grouted tendons, there have been 

sufficient corrosion problems in the non-nuclear industries using grouted tendons to warrant the 

need for improved monitoring methods of grouted tendons. Because the prestressing level in 

grouted tendons cannot be measured directly, the use of non-destructive testing methods and 

sensing technologies appear to be a practical way to monitor the prestressing system when the 

tendons are grouted. Countries around the world are using sensors to monitor grouted tendons 

for nuclear applications. In addition the development of reliable NDE techniques will lead to 

increased confidence for monitoring. With the new construction of EPRs with grouted tendons, 

valuable experience and information will be gained from Olkiluoto 3 in Finland, Flamanville 3 in 

France, and planned construction of the UKEPR, and MDEP.

5.1 Comparison of Structural Response of Grouted and Ungrouted 

Tendon Systems

Modeling techniques used in this work build on those developed for International Standard 

Problem 48 and represent significant advances. To improve computational efficiency, 

particularly for the global 3D model, the use of slot connectors with friction in Abaqus have been 

implemented for ungrouted tendons and have been shown to be sufficiently accurate when 

compared to the contact surface approach. The slot connector approach captures the initial 

stress state with respect to position along the length of the tendons including the losses 

associated with friction and with anchoring. This is an important improvement as the 

computational efficiency of this approach facilitates global 3D modeling of the entire 

containment with the individual tendons represented. The level of detail represented in this 

global model has not been achieved in any previous modeling effort. With this approach, 

globally significant local effects are captured such as the changing tendon stress distribution 

with internal pressurization. Modeling the ungrouted tendons in this way leads naturally to 

modeling grouted tendons. Then, the effect of bonding the tendons to the concrete structure is 

captured with minimal computational expense by exchanging the slot connectors for rigid beam 

connectors. Additionally, the Abaqus concrete damaged plasticity model has been successfully 

implemented thus eliminating the need for third party concrete models (e.g. Anatech Anacap 

model) which add additional computational expense.

The analytical work performed and documented in this report, have led to the following 

conclusions regarding the structural behavior of grouted and ungrouted prestressing systems for 

concrete containment buildings. In general, for PCCVs, when comparing grouted to ungrouted

tendon behavior, cylinder deformations are larger with grouted tendons, all other things being 

equal. This appears to be related to the inability of the grouted tendons to slip and redistribute 

load. Accordingly, at high pressures, tendon strains (and forces) are less well distributed along 

the length of the tendons for grouted tendon systems. This is because local increments of strain 

in the vessel wall must track one-for-one with tendon strain increments; not so for ungrouted
5-2

tendons. Accordingly, maximum strains are larger with grouted tendons versus ungrouted 

tendons for the un-corroded case. For these reasons, it is expected that PCCVs with grouted 

prestressing systems will have a lower ultimate pressure capacity than PCCVs with ungrouted 

prestressing systems based either on leakage or structural failure owing to the greater localized 

deformation and strain which would lead to containment failure. 

Unfortunately, the experimental basis for evaluating containment performance with grouted 

prestressing systems is limited. A recently completed containment test in India [Parmar, Singh 

et al. 2014] included a grouted prestressing system. Additionally planned containment tests in 

France will also include grouted tendons. Both tests should provide useful data to further 

evaluate the model predictions presented here.

5.2 Grouted and Ungrouted Tendon System Comparison

A review of standards for grouted and ungrouted tendon systems has been organized and 

presented in section 3.1. Similarly, international nuclear industry regulations and practices have 

been summarized in section 3.2.2.3. These sections serve as useful references for comparing 

existing knowledge on the topic of grouted and ungrouted prestressing systems. 

Grouted and ungrouted tendons systems each have advantages and disadvantages that have 

been investigated in this research. The most significant points of comparison are summarized 

below:

• Corrosion protection: The prestressing steel is protected from corrosion with grease in 

an ungrouted system and with cementitious grout in a grouted system. The grease coats 

the tendon steel and thus prevents oxygen, moisture, and deleterious ions from reaching 

the steel while cementitious grout provides a barrier through the high alkalinity of the 

grout which passivates the steel tendon. 

• Tendon interaction with the concrete structure: The interaction between the tendon and 

the concrete structure is fundamentally identical between the two systems up to the point 

of grouting. After grouting, and in response to applied loading, the behavior of the two 

systems differs in the ability of the tendon to move relative to the structure. For the 

ungrouted system, load transfer between the tendons and the reinforced concrete 

structure occurs through the anchorages, through normal forces and through frictional 

forces. Grouted systems transfer load continuously through the grouting. Grouting also 

has a beneficial effect of providing redundancy to the anchorage since the length of the 

tendon becomes “anchored” by the grout. Similarly, a local defect in the tendon remains 

local, i.e. the tendon force is not affected over the entire tendon length. In the context of 

beyond design basis type loading, the present modeling presented in Sections 2.3.3 and 

2.4.2 indicates that this localization may actually reduce the structural performance since 

the localized deformations tend to be larger which leads to failure. On the other hand, 

when significant tendon corrosion is considered, performance of grouted systems 

appears to be better as documented in Sections 4.8 and 4.10. 

• Future adjustment of prestressing force: Prestressing forces of ungrouted tendons can 

be adjusted at any time during the design life of a structure. However, all necessary 

tendon details for later stressing need to initially be provided such as access and 

clearance for jacks, and sufficient extra length of prestressing steel to connect the jack to 

the strand. This is considered impossible for grouted systems.
5-3

• Tendon replacement: Similar to above, the ability to replace damaged tendons is 

afforded with ungrouted prestressing systems and is not with grouted prestressing 

systems. Furthermore, the ability to remove and replace tendons can facilitate large 

system repairs to the plant by opening the containment building after the removal of 

prestress. 

• Facilitated inspection of tendon: Inspecting the prestressing tendons is much more

straightforward with the ungrouted system. For grouted systems access to the tendon is 

limited to sensors, NDE techniques, or targeted invasive inspection. Hence, special

inspection or monitoring devices still need to be used to collect information on the actual 

performance and durability of the steel for grouted systems.

5.3 Tendon Corrosion for Grouted and Ungrouted Systems

An analytical method involving reduction of tendon area provides a reasonable approach to 

simulating corrosion, and solutions are obtainable. Expansion of the tendon due to corrosion 

was not considered nor was the associated cracking of grout and concrete. The loss of bond 

with the tendon was also not considered. It was assumed that the effect of corrosion on the load 

capacity of the tendon was the dominant mechanism affecting the ultimate load capacity of the 

containment and that the aforementioned phenomena were secondary. Reduction of area of 

about 60% is the apparent limit (based on this short study) for obtaining solutions in this way. 

Area reductions more than 60% failed to converge when only subjected to the prestressing 

forces (e.g. the structure has difficulty redistributing the prestress when 60% is lost along a 

substantial bank of tendons). Liner tearing and leakage were not addressed in this study, rather 

the structural performance and viability of the system in response to internal pressurization was 

studied. Furthermore, only internal pressurization was considered and conclusions presented 

should be considered representative of other loading cases (e.g. seismic).

In general, grouted prestressing systems appear to have a significant advantage for the 

corrosion cases studied. For each of these studied corrosion cases, the ungrouted systems 

demonstrated significantly reduced ultimate structural capacity while the grouted systems 

showed a smaller reduction in ultimate capacity when compared to the uncorroded 

configuration. 

When vertical tendon corrosion was studied (Case 1), ungrouted simulations resulted in a 

reduction in ultimate pressure capacity (tendon failure capacity, NOT capacity defined by liner 

tearing) of about 10% (3.3Pd versus 3.6Pd). For the grouted tendon system the ultimate 

pressure capacity is nearly the same as for the corresponding uncorroded case. The main 

difference between the uncorroded and corroded case is the strains in the corroded cases are 

slightly higher.

For hoop tendon corrosion near the E/H (Case 2), the ungrouted simulations resulted in 

reduction in ultimate pressure capacity of approximately 25%, with tendons beginning to rupture 

between 2.3 and 2.5Pd, and pressure unsustainable beyond 2.8Pd. The reduction in capacity 

would be even greater if tendon embrittlement (reduction in ductility) were considered. The 

grouted hoop tendon simulations with corrosion indicated larger maximum tendon strain, but 

similar ultimate pressure capacity based on tendon behavior to the uncorroded, grouted 

simulation. 

Combining the vertical tendon corrosion and hoop tendon corrosion, the ungrouted simulation 

results were similar to hoop tendon corrosion case, except with tendons failing between 2.0 and 

2.3Pd, and pressure unsustainable beyond 2.3Pd which would correspond to an approximate 
5-4

reduction in capacity of 35%. As before, the corroded grouted simulations were comparable to 

the uncorroded, grouted simulation with the maximum tendon strains being slightly higher and 

the ultimate pressure capacity being similar.

The hoop tendons at the anchor zone corrosion case, ungrouted resulted in ultimate pressure 

capacity reduction of approximately 10%, with tendons beginning to rupture at 2.5Pd, and 

pressure unsustainable beyond 3.2Pd based on tendon behavior. The corroded, grouted 

simulations resulted in similar performance to the uncorroded simulation.

From the simulations summarized here and in agreement with pressure vessel theory, the most 

critical area of corrosion that causes early failure of the PCCV for both ungrouted and grouted 

cases occurs at the hoop tendons near the E/H. 

The apparently superior performance of grouted prestressing systems in the context of tendon 

corrosion can be explained in a few steps. As a given tendon corrodes, the force in the tendon 

decreases due to increased deformation associated with loss of effective area. This has the 

general effect of allowing increased radial displacement of the cylinder wall, and increased liner 

and rebar strains (increased hoop tendon strains as well). Grouting the tendons makes the force 

next to the corrosion decrease more, as shown in Figure 4-36. As shown earlier, however, the 

extent of tendon force loss is very localized near the corrosion, whereas an ungrouted tendon 

has a larger force next to the corroded anchor but distributes the loss over a wider region. 

Stated more succinctly, the ungrouted system loses prestressing over a much greater region 

than the grouted system, for a given level of corrosion, and this large loss of prestress results in 

greater deformations which ultimately drive failure. This is particularly true in areas of the vessel 

with higher stresses from internal pressurization such as the hoop tendons near the equipment 

hatch. It is interesting to note that the same localizing phenomenon resulted in poorer 

performance when the prestressing system was uncorroded. While tendon corrosion shouldn’t 

be specifically planned for, the extra capacity afforded by grouting the tendons in the case of 

tendon corrosion should be considered an advantage of grouted tendons.

5.4 Overall Conclusions

The question of grouting prestressing tendons in nuclear power plant containment structures 

raises many questions and some of these are addressed here. The modeling presented here 

was based on the 1:4 Scale PCCV Test geometry and the associated data and experience 

gained in previous modeling efforts for the 1:4 Scale PCCV. Unfortunately, no analogous test 

data is available for a grouted tendon system, so the modeling predictions associated with the 

grouted tendons are only based on extensive analyst expertise, but not benchmarked to test 

data. These predictions indicate a slight reduction in ultimate pressure capacity for grouted 

tendons compared to ungrouted tendons when the tendons are uncorroded, and conversely a 

significant advantage for grouted tendons compared to ungrouted tendons when significant 

corrosion of the tendons is present. 

Grouted prestressing systems are significantly more difficult to monitor over the life of the 

structure owing to the reduced access to the tendons themselves. NDE techniques or global 

proof testing (e.g. internal pressure testing) are options for monitoring grouted tendons. For 

ungrouted tendons, more options exist for monitoring and inspection including lift-off testing and 

physical removal of individual tendons for inspection. Furthermore, the ability to remove 

tendons provides options for responding to unanticipated structural degradation such as tendon 

corrosion and facilitates internal repair activities such as steam generator replacement where 

the containment needs to be opened. 
5-5

Global experience with grouted prestressing systems in reactor containment buildings will be 

enhanced in the coming years as EPRs are built and are operated. This experience will be 

valuable in regulatory decision making, but the central question of grouted versus ungrouted 

tendons appears to be tied to the relative importance of inspection and monitoring in 

comparison to improved structural performance. Ungrouted tendon systems have the unique 

advantage of post construction modification, which is useful for inspection, repair, and 

adjustment.

6-1

6 REFERENCES

AASHTO LRFD Bridge 2013, "AASHTO LRFD Bridge Design Specifications," AASHTO, 

Washington, DC, 2013.

"Abaqus 6.11" (2011). Dassault Systèmes Simulia Corp., Providence, RI.

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NUREG/CR-7208 Study on Post Tensioning Methods