Structural safety of industrial steel tanks, pressure vessels and piping systems under seismic loading (INDUSE).
Date: 2014-01-15
Institutions: Directorate-General for Research and Innovation
Subjects: environmental risk prevention, industrial infrastructure, industrial research, iron and steel industry, iron and steel product, natural disaster, piping, pressure equipment, research report, resistance of materials, steel

European Commission

Research Fund for Coal and Steel

Structural safety of industrial steel tanks,  
pressure vessels and piping systems  
under seismic loading
(INDUSE)

P. Pappa, G. E. Varelis, M. Vathi, P. C. Perdikaris, S. A. Karamanos
University of Thessaly
Argonafton & Filellinon, 38221 Volos, GREECE
J. Ferino, A. Lucci, E. Mecozzi, G. Demofonti
Centro Sviluppo Materiali S.p.A.
Via di Castel Romano, 100/102, 00128, ITALY
A. M. Gresnigt, G. J. Dijkstra (Tebodin)
Delft University of Technology
Stevinweg 1, 2628 CN Delft, NETHERLANDS
Md. S. Reza, A. Kumar, F. Paolacci, O. S. Bursi
University of Trento
Via Belenzani 12, 38100 Trento, ITALY
M. Kopp, M. Pinkawa, M. Wieschollek, B. Hoffmeister
RWTH Aachen University
Templergraben 55, 52062 Aachen, GERMANY
A. Stamou, K. Diamanti, T. Papatheocharis
EBETAM (MIRTEC) S.A.
A’ Viomichaniki Periochi, 38500 Volos, GREECE
C. Botsis, I. Chandrinos, I. Doukas
TECHNIPETROL HELLAS S.A.
Pavlou Bakogianni Street 42, 14452 Metamorfosis, Athens, GREECE

Grant Agreement RFSR-CT-2009-00022 
1 July 2009 to 31 December 2012

Final report

Directorate-General for Research and Innovation

2013 

EUR 26319 EN

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Luxembourg: Publications Office of the European Union, 2013 

ISBN 978-92-79-34575-3 
doi:10.2777/49423 

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Table of contents 

Table of contents ....................................................................................................................................3 

I. 

Final summary ..............................................................................................................................7 

I.1.  WP1:  Comparison  of  European  and  American  design  standards  for  the  seismic  design  of 
industrial  liquid  storage  tanks,  pressure  vessels  and  piping  and  development  of  available  test  data 
base. 

 ..................................................................................................................................................8 

I.2.  WP2:  Determination  of  seismic  forces  on  industrial  structures  and  critical  components  and 
preparation of experiments. ..................................................................................................................8 

I.3.  WP3: Experimental testing of structural components of critical importance under strong cyclic 
loading. .................................................................................................................................................9 

I.4.  WP4: Numerical simulation of industrial component behaviour under strong cyclic loading.10 

I.5.  WP5:  Seismic  behaviour  and  special  seismic  design  issues  in  liquid  storage  tanks; 
development of relevant design guidelines/recommendations. ..........................................................12 

I.6.  WP6:  Seismic  behaviour  and  design  of  industrial  pressure  vessels;  development  of  seismic 
design guidelines/recommendations. ..................................................................................................13 

I.7.  WP7:  Seismic  behaviour  and  design  of  industrial  piping  systems  and  components; 
development of seismic design guidelines/recommendations. ...........................................................14 

I.8.  WP8: Coordination of the project. ..........................................................................................15 

I.9. 

Conclusions and main results .................................................................................................15 

II. 

Scientific description of the results ............................................................................................17 

Objectives of the project .....................................................................................................................17 

Comparison of initially planned activities and work accomplished ...................................................17 

II.1.  Work  package  1  –  Comparison  of  European and  American  design  standards  for  the  seismic 
design  of  industrial  liquid  storage  tanks,  pressure  vessels  and  piping  and  development  of  available 
test data base .......................................................................................................................................17 

II.1.1. 

Task 1.1: Collection and evaluation of available relevant experimental data ................17 

II.1.2. 
facilities   ........................................................................................................................................18 

Task  1.2:  Collection  and  evaluation  of  reported  earthquake  damages  in  industrial 

II.1.3. 

Task 1.3: Basic seismic structural design of typical liquid storage tanks.......................19 

II.1.4. 
vessels 

Task  1.4:  Basic  seismic  structural  design  of  typical  cylindrical  and  spherical  pressure 
 ........................................................................................................................................19 

II.1.5. 
structure   ........................................................................................................................................20 

Task  1.5:  Basic  seismic  structural  design  of  typical  piping  system  and  its  support 

II.2.  Work  package  2  –  Determination  of  seismic  forces  on  industrial  structures  and  critical 
components and preparation of experiments ......................................................................................22 

II.2.1. 

Task 2.1: Seismic forces in vertical cylindrical liquid storage tanks .............................22 

II.2.2. 

Task 2.2: Seismic forces in pressure vessels ..................................................................22 

II.2.3. 

Task 2.3: Seismic forces in industrial piping systems ....................................................23 

II.2.4. 

Task 2.4: Dissipative support systems ............................................................................24 

II.2.5. 

Task 2.5 Determination of loading amplitudes on critical components .........................24 

II.2.6. 
supply 

Task  2.6  Specification  of  procedures  for  specimen  manufacturing  and  base  material 
 ........................................................................................................................................24 

II.3.  Work package 3 – Experimental testing of structural components of critical importance under 
strong cyclic loading ..........................................................................................................................25 

II.3.1. 

Task 3.1: Basic material testing and characterization ....................................................26 

3 
II.3.2. 

Task 3.2: Tests on tank openings/nozzles under strong cyclic loading ..........................27 

II.3.3. 

Task 3.3: Tests on bolted piping connections under strong cyclic loading ....................29 

II.3.4. 

Task 3.4: Tests on piping elbows under strong cyclic loading .......................................31 

II.3.5. 

Task 3.5: Tests on T-branch pipe connections under strong cyclic loading ...................35 

II.4.  Work  package  4  –  Numerical  simulation  of  industrial  components  behaviour  under  strong 
cyclic loading .....................................................................................................................................37 

II.4.1. 

Task 4.1 Definition of inelastic material modelling under cyclic loading .....................37 

II.4.2. 
importance, tested in WP3 ..............................................................................................................41 

Task  4.2  Numerical  simulation  of  the  experimental  tests  on  components  of  critical 

II.4.3. 

Task 4.3 Parametric numerical study of critical component behaviour .........................55 

II.5.  Work package 5 – Seismic behaviour and special seismic design issues in liquid storage tanks; 
development of relevant design guidelines/recommendations ...........................................................65 

II.5.1. 

Task 5.1 Special issues in seismic behaviour of liquid storage tanks.............................65 

II.5.2. 
liquid storage tanks .........................................................................................................................68 

Task  5.2  Design  recommendations  /  guidelines  for  the  seismic  analysis  &  design  of 

II.6.  Work package 6 –  Seismic behaviour and design of industrial pressure vessels; development 
of seismic design guidelines/recommendations .................................................................................69 

II.6.1. 

Task 6.1 Seismic behaviour of vertical-cylindrical pressure vessels..............................69 

II.6.2. 

Task 6.2 Seismic behaviour of horizontal-cylindrical pressure vessels .........................69 

II.6.3. 

Task 6.3 Seismic behaviour of pressure vessels of spherical shape ...............................70 

II.6.4. 
analysis of industrial pressure vessels ............................................................................................70 

Task 6.4 Development of design recommendations / guidelines for the seismic design & 

II.7.  Work  package  7  –  Seismic  behaviour  and  design  of  industrial  piping  systems  and 
components; development of seismic design guidelines/recommendations ......................................71 

II.7.1. 

Task 7.1 Seismic behaviour of industrial piping and their support systems...................71 

II.7.2. 
& design of industrial piping systems ............................................................................................72 

Task 7.2 Development of design recommendations / guidelines for the seismic analysis 

II.8.  Work package 8 – Coordination of the project.......................................................................73 

II.8.1. 

Task 8.1 Efficient management process .........................................................................73 

II.8.2. 

Task 8.2 Definition of the standard protocols ................................................................74 

II.8.3. 

Task 8.3 Monitoring of project activities and evaluation of the project .........................74 

II.8.4. 

Task 8.4 Progress, Mid-Term and Final reports .............................................................75 

II.9.  Conclusions ............................................................................................................................76 

II.10. 

Exploitation and impact of the research results ..................................................................76 

III.  Appendices .................................................................................................................................77 

III.1. 

Appendix of WP1 ...............................................................................................................77 

III.1.1.  Task 1.1: Collection and evaluation of available relevant experimental data ................77 

III.1.2.  Task 1.3: Basic seismic structural design of typical liquid storage tanks.......................77 

III.1.3.  Task  1.4:  Basic  seismic  structural  design  of  typical  cylindrical  and  spherical  pressure 
vessels 

 ........................................................................................................................................78 

III.1.4.  Task  1.5:  Basic  seismic  structural  design  of  typical  piping  system  and  its  support 
structure   ........................................................................................................................................78 

III.2. 

Appendix of WP2 ...............................................................................................................79 

III.2.1.  Task 2.1: Seismic forces in vertical cylindrical liquid storage tanks .............................79 

4 
III.2.2.  Task 2.2: Seismic forces in pressure vessels ..................................................................80 

III.3. 

Appendix of WP3 ...............................................................................................................81 

III.3.1.  Task 3.1 Basic material testing and characterization .....................................................81 

III.3.2.  Task 3.2: Tests on tank openings/nozzles under strong cyclic loading ..........................85 

III.3.3.  Task 3.3: Tests on bolted piping connections under strong cyclic loading ....................88 

III.3.4.  Task 3.4: Tests on piping elbows under strong cyclic loading .......................................94 

III.3.5.  Task 3.5: Tests on T-joint pipe connections under strong cyclic loading ....................103 

III.4. 

Appendix of WP4 .............................................................................................................106 

III.4.1.  Task 4.3 Parametric numerical study of critical component behaviour .......................106 

III.5. 

Appendix of WP5 .............................................................................................................113 

III.5.1.  SEISMIC RESPONSE OF LIQUID STORAGE TANKS ...........................................113 

III.6. 

Appendix of WP6 .............................................................................................................116 

III.6.1.  Task 6.1 Seismic behaviour of vertical-cylindrical pressure vessels............................116 

III.6.2.  Task 6.2 Seismic behaviour of horizontal-cylindrical pressure vessels .......................118 

III.6.3.  Task 6.3 Seismic behaviour of pressure vessels of spherical shape .............................120 

III.7. 

Appendix of WP7 .............................................................................................................124 

III.7.1.  General analyses procedure for seismic design acc to EN 1998-1 and 4 .....................124 

III.7.2.  Seismic linear elastic analyses methods .......................................................................125 

III.7.3. 

Investigation of seismic input .......................................................................................125 

III.7.4. 

Influence of supporting structure ..................................................................................127 

III.7.5.  Example of ultimate limit state for elbows ...................................................................128 

IV.  List of acronyms and abbreviations ..........................................................................................131 

V.  List of figures ...........................................................................................................................133 

VI.  List of tables .............................................................................................................................139 

VII.  References ................................................................................................................................141 

5 
 
 
I.  Final summary 

Liquid  storage  tanks,  pressure  vessels  and industrial piping  are  steel  structural systems,  present  in  all 
industrial  facilities  (chemical  and  petrochemical  industries,  and  power  plants).  Safeguarding  their 
structural integrity against earthquakes constitutes a key issue towards increased safety and unhindered 
operation of the industrial facility. Their structural behaviour, and in particular seismic design, is quite 
different than steel buildings and has several particularities, requiring a combined civil and mechanical 
engineering  expertise.  The  particularities  stem  from  their  shape  and  geometry,  the  presence  of  high 
internal pressure, and the dynamic behaviour of the enclosed liquid, which may affect significantly their 
load and deformation capacity.  

The design of those structures has been dominated by the use of American standards (API 650, ASME 
VIII  and  B31.3).  Nevertheless,  for  the  case  of  seismic  design,  those  standards  (especially  ASME 
standards for pressure vessels and piping) contain very limited provisions, referring mainly to structural 
design  codes  and  specifications.  European  specifications  (EN  14015,  13445,  13480)  also  contain 
limited provisions for the earthquake-resistant design of industrial equipment. An effort has been made 
in Eurocode 8 (EN 1998-4), which concerns almost exclusively vertical-cylindrical liquid storage tanks. 
Nevertheless,  those rules do not cover all possible limit states and, furthermore, they need significant 
improvement to reach a level of applicability for design practice.  

INDUSE

The 
  program  consists  of  an  interdisciplinary  research  effort  that  combines  civil  and 
mechanical engineering expertise, for the purpose of developing guidelines, which can be used for the 
seismic  design  of  liquid  storage  tanks,  pressure  vessels  and  piping,  within  the  Eurocode  design 
framework.  Towards  this  purpose,  extensive  experimental,  analytical,  and  numerical  work  has  been 
conducted  within  the 
  project  with  the  synergy  of  academic  units,  research  centres  and 
industrial partners. The guidelines are novel and unique, incorporating modern aspects of earthquake-
resistant design, and are aimed at: 

INDUSE

  expanding EN 1998-4 provisions towards an integrated seismic design of liquid storage tanks 
and  attached  equipment,  incorporating  some  special  features  and  all  possible  failure  modes 
(WP5) and 

  extending  the  applicability  of  Eurocode  8  (EN  1998)  concepts  for  the  cases  of  industrial 

pressure vessels (WP6), and piping systems (WP7). 

The following intermediate goals have been achieved within the 

INDUSE

 project: 

  A  basic  comparison  has  been  performed  between  current  seismic  design  provisions  in 
European  and  American  standards  together  with  an  assessment  of  seismic  damages  in 
industrial facilities (WP1).  

  Taking  into  account  the  particularities  of  each  structural  system,  seismic  actions  have  been 

determined (WP2). 

  Extensive experimental testing has been conducted on key piping components (e.g. nozzles, 
pipe  connections/branches  and  elbows)  under  strong  cyclic  loading;  furthermore  a  piping 
system has been tested under pseudo-dynamic loading (WP3).  

  Finite  element  analyses  on  these  components  have  been  performed,  simulating  the 
experiments and covering a wide range of geometric, material and loading parameters (WP4). 

The  results  of  the  above  investigations  are  summarised  below,  demonstrating  the  achievements  of 
INDUSE

 project objectives, work package per work package. 

7 
 
 
I.1.  WP1:  Comparison  of  European  and  American  design  standards  for  the 
seismic design of industrial liquid storage tanks, pressure vessels and piping and 
development of available test data base. 

The objectives of this workpackage consist of (a) collecting and evaluating available cyclic loading test 
data  on  critical  industrial  components,  (b)  collecting  and  evaluating  damages  in  industrial  equipment 
due  to  strong  earthquakes  and  (c)  comparing  relevant  European  and  American  standards  in  terms  of 
seismic  design.  In  task  1.1,  the  following  experimental  testing  programs  on  piping  components  and 
piping systems are described, evaluated and extensively discussed: 

Berkeley National Laboratory tests, UK  
University of Liverpool tests (1st and 2nd part) 
EPRI-NRC Piping and Fitting Dynamic Reliability Program 
ETEC System Tests 

1. 
2. 
3. 
4. 
5.  Westinghouse Hanford Test   
6. 
EPRI Prototype Tests 
7. 
Japanese High Level Tests 
8. 
University of Akron Studies 
9. 
Japanese Vibration Tests 

The collection and evaluation of typical damages in industrial equipment structures subjected to strong 
earthquake events has been performed in task 1.2, for the following earthquake events: 

1. 
2. 
3. 
4. 
5. 
6. 

Imperial Valley Earthquake, California, USA, October 15, 1979 
Coalinga Earthquake, California, USA, May 2, 1983 
Northridge Earthquake, Los Angeles, USA, January 17, 1994 
Kocaeli Earthquake, Turkey, August 17, 1999 
Chile Earthquake, March 3, 1985 
Kobe Earthquake, January 17, 1995 

Finally,  the  main  effort  of  WP1  has  concentrated  in  task  1.3,  on  the  selection  and  preparation  of  the 
seismic design for eleven (11) case-studies on tanks, pressure vessels and piping, namely: 

 
 
 
 
 
 
 
 
 
 
 

Case Study #1: 27.4-m-diameter liquid storage tank  
Case Study #2a: 18-m-diameter liquid storage tank  
Case Study #2b: 6-m-diameter liquid storage tank  
Case Study #2c: liquid storage tank with floating roof  
Case Study #3: broad liquid storage tank  
Case Study #4: spherical pressure vessel  
Case Study #5: vertical pressure vessel 
Case Study #6: horizontal pressure vessel  
Case Study #7: vertical pressure vessel on a steel structure  
Case Study #8: piping system on a pipe-rack  
Case Study #9: piping system on a steel support structure  

The  above  case-studies  refer  to  existing  industrial  structures,  and  were  provided  by  our  industrial 
partners.  Their  seismic  design  has  been  performed  using  European  and  American  standards.  Seismic 
actions have been calculated, and the necessary verification checks were conducted. Inconsistencies and 
issues that require further consideration have also been identified. 

I.2.  WP2:  Determination  of  seismic  forces  on  industrial  structures  and  critical 

components and preparation of experiments. 

The objectives of this workpackage (WP2) have been: (a) the development of systematic procedures to 
calculate  seismic  action  on  tanks,  pressure  vessels  and  piping,  using  advanced  analysis  methods, 
considering the particularities of the above systems; (b) investigation of the effects of nonlinear support 
system response on the seismic behaviour of pressure vessels and piping; examination of the influence 
of  dissipative  support  systems  towards  seismic  force  reduction;  (c)  Determination  of  cyclic  loading 
earthquake action on selected critical components, considering strong seismic input, and employing the 

8 
 
 
above  analysis  procedures;  (d)  Specification  of  specimen  geometry  to  be  tested  in  the  subsequent 
experimental program together with the detailed specification of base material and welding procedures. 

The above work has been accomplished in the course of WP2. In particular,  

  In  Tasks  2.1,  2.2,  and  2.3,  procedures  to  calculate  seismic  action  on  industrial  structures  have 
been presented, using advanced analysis methods, for liquid storage tanks, pressure vessels and 
piping, respectively. The methodologies are based on advanced numerical methods, taking into 
account the dynamic interaction between the structure and the enclosed liquid (in liquid storage 
tanks) and the interaction between the piping and the supporting structure (in piping systems). 
  The  investigation  of  nonlinear  support  system  response  on  the  seismic  behaviour  of  pressure 

vessels, for the purpose of reducing of the seismic force, has been investigated in Task 2.4.  

  A typical piping system has been considered, and simulated through advanced numerical tools in 

order to determine seismic action in various components, as described in Task 2.5.  

  Finally, in Task 2.6 specifications for the ordering and manufacturing of specimens to be tested 
in  the  subsequent  experimental  program  have  been  specified.  Detailed  specifications  of  base 
material and welding procedures have been stated to be followed by all partners. 

I.3.  WP3:  Experimental  testing  of  structural  components  of  critical  importance 

under strong cyclic loading. 

In  this  workpackage,  extensive  experimental  testing  has  been  performed  on  steel  material,  on  piping 
components and on a piping system. 

Testing  has  been  performed  for  the  mechanical  characterization  of  both  base  material  and  weld 
material.  Base  material  characteristics  has  been  evaluated  for  elbows;  tee-fittings;  plates  for  nozzles; 
straight pipes connected to fittings in full-scale specimens. The base material characterization consists 
of  (a)  monotonic  tensile  testing;  (b)  cyclic  loading  with  imposed  strains  to  obtain  hysteresis  loops 
(cyclic  hardening/softening)  of  elbows  both  longitudinally  and  transversely;  (c)  cyclic  loading  with 
imposed stresses to obtain ratcheting characteristics of elbows both longitudinally and transversely; (d) 
cyclic loading with imposed deformations to obtain cyclic hardening/ softening of straight pipes, to be 
connected weld neck flanges in full scale testing; (e) ring crush testing to evaluate the possible material 
anisotropy of straight pipes, to be connected to flanges and elbows in full scale testing. In addition, tests 
on  butt  welded  specimens  in  8-inch  pipes,  to  check  the  validity  of  the  WPS  (developed  within  the 
INDUSE
 project). First, non-destructive tests were carried out. Then, destructive tests were performed: 
(a) macroscopic examination of welds (EN1321); (b) hardness tests (EN1043-1); (c) bend test on welds 
(EN910); (d) transverse tensile test on welds (EN895, EN6892-1); (e) Charpy V-notch impact test on 
welds (EN875, EN10045-1). Low cycle fatigue tests of welded specimens were also carried out and 
an S-N curve was derived. 

Subsequently,  a  number  of  medium-scale  experiments  in  key  components  of  industrial  structural 
systems have been conducted under strong cyclic loading, to determine their mechanical behaviour and 
failure modes, and to obtain adequate measurements for the calibration of the numerical models, to be 
used in subsequent work packages. The components are 6-inch and 8-inch pipes, and the material of all 
specimens  is  P355N,  which  is  equivalent  to  X52.  In  addition  to  the  tests  on  specific  components,  a 
piping system, containing 3 elbows, 1 Tee junction and 1 flange, has been tested  tests under  pseudo-
dynamic loading at 4 seismic levels. More specifically,  

  six  (6)  tests  have  been  conducted  on  6-inch  tank  openings  (nozzles).  Three  different  shell  nozzle 
reinforcements are considered, namely (P1) internal reinforcing plate, (P2) external reinforcing plate 
and (P3) thickened nozzle reinforcement. The tests were carried out in the range of ultra-low-cycle 
fatigue considering longitudinal and transverse load directions for each nozzle reinforcement type. 
Increasing  amplitude  loading  was  applied  according  to  ECCS  No.45  protocol.  Failure  occurred  in 
form of fatigue crack in either side of the weld between the 6-inch pipe and the reinforced shell. All 
three  types  of  reinforcement  exhibited  similar  strength  in  terms  of  loading  cycles.  Furthermore, 
transverse loading has been more severe than longitudinal loading.  

  fifteen (15) tests have been conducted on full-scale bolted (flanged) piping connections as follows: 
(a)  Six  tests  have  been  conducted  on  standard  PN40  and  PN63  bolted  flanged  connections,  as 

9 
 
specified  in  EN  1092,  subjected  to  monotonic  and  cyclic  bending  load  conditions,  in  presence  of 
internal pressure. (b) An experimental test program on eight (8) non-standard Bolted Flange Joints 
has  also  been  conducted  under  monotonic  and  cyclic  loading  with  increasing  amplitude  (in  the 
presence of low internal pressure) in order to investigate their capacity and performance and assess 
their  capacity  under  seismic  loading.  Two  non-standard  bolted  connection  geometries  have  been 
considered, comparatively thinner than the Standard ones, connected with an 8-inch connecting pipe 
under regular operating conditions of a petrochemical industry.  

  A  significant  number  of  tests  [thirteen  (13)  tests]  have  been  performed  on  8-inch  (D=219.1mm) 
long-radius  (R/D=1.5)  SCH40  (t=8.2mm)  elbow  specimens,  from  P355N  material,  connected  to 
straight  pipes  of  the  same  dimensions  at  either  side  (L/D=5),  under  constant-amplitude  in-plane 
loading. Eight non-pressurized specimens have been tested at various loading amplitudes (covering 
the entire range from quasi-elastic to severely-plastic behaviour). In addition, five specimens have 
been tested under pressure (at pressure levels up to 42% of yield pressure) at strong cyclic loading, 
causing severe plastic deformation. In all specimens, failure occurred in the form of a longitudinal 
crack at the central part of the elbow, at the flank location. It was found that the number of cycles to 
failure  depends  significantly  on  the  loading  amplitude,  whereas  the  presence  of  internal  pressure 
resulted in a small decrease of the number of loading cycles to failure.  

  Eight  (8)  tests  have  been  performed  on  8-inch-to-6-inch  SCH40  Tee-branch  pipe  connections, 
considering two configurations: (a) four specimens with Tee fittings (TJF) and (b) four specimens 
with reinforcement plates at the connection of its two main parts (TJWR). The specimens have been 
tested  under  monotonic  loading  (1  test  for  each  type)  and  cyclic  out-of-plane  bending  loading  (3 
tests  for  each  type).  Under  monotonic  loading  conditions,  the  TJF  specimen  showed  about  93% 
higher  load  capacity.  Under  cyclic  loading  conditions,  the  TJF  specimen  exhibited  higher  fatigue 
resistance of about 7-8 times that of the TJWR specimen. For the increasing amplitude loading, the 
TJF specimen resisted 36% more loading cycles than the TJWR one. The TJWR specimens failed 
along the circumference of the branch-to-reinforcement plate weld-toe either on the branch side or 
on the welded reinforcement plate side of the connection. The TJF specimens failed in the region of 
the joint fitting, at a distance of about 60 to 80 mm away from the branch-to-fitting weld-toe. 

Finally, in addition to testing individual components under strong cyclic loading, a piping system has 
been tested through pseudo-dynamic loading, at full-scale. The piping system contained several critical 
components,  mainly  elbows,  also  1  flange,  and  1  Tee,  and  it  is  aimed  at  investigating  its  seismic 
performance at four seismic levels, offering a good alternative to more expensive shaking table tests. 
The  present  testing  employs  a  new  hybrid  testing  technique  which  enables  the  seismic  testing  of  a 
structural system by using a limited number of actuators. Several pseudo-dynamic tests on the piping 
system were performed under different levels of earthquake PGAs corresponding to serviceability and 
ultimate  limit  states  suggested  by  Standards.  The  experimental  and  numerical  results  showed  a  good 
agreement in all tests. Moreover, a good seismic performance of the piping system was observed during 
the  experiments.  It  was  found  that,  even  under  the  Collapse  Limit  State  level  earthquake,  both  the 
piping system and its critical components did not fail or yield. Therefore, it is concluded that present 
seismic  design  rules  for  piping  systems  and  components  are  rather  conservative  and  amendments  are 
necessary to overcome some degree of conservatism. 

I.4.  WP4:  Numerical  simulation  of  industrial  component  behaviour  under  strong 

cyclic loading. 

The  main  purpose  of  this  work  package  is  the  development  of  an  advanced  numerical  simulation 
procedure, and its use for the prediction of mechanical behaviour of critical structural components, in a 
rigorous, efficient and accurate manner. The specific objectives of this workpackage (WP4) have been: 
(a) the definition and calibration of an appropriate cyclic-loading constitutive material model to use for 
a realistic simulation of elasto-plastic behaviour of tubular components; the constitutive material model 
should  be  able  to  simulate  material  degradation  effects,  as  well  as  low-cycle  fatigue,  (b)  the 
development of a modelling procedure  using finite element that simulates accurately the behaviour of 
industrial components; the modelling procedure has been validated through the experimental data of the 
previous  workpackage  (WP3)  and  (c)  the  performance  of  an  extensive  parametric  study,  using  the 
numerical  modelling  methodology  developed  above,  to  examine  the  behaviour  of  structural 
components, within a wide range of geometric and material parameters. 

10 
 
INDUSE

Task  4.1  has  been  devoted  on  the  choice,  implementation  and  calibration  of  a  suitable  material 
(constitutive)  model  for  simulating  the  material  behaviour  of  steel  material.  Following  significant 
investigation, it has been concluded that the classical plasticity models based on isotropic and kinematic 
hardening rules may not be adequate in simulating cyclic response of steel material under both strain-
controlled  and  stress-controlled  conditions.  Under  this  perspective,  the  Armstrong-Frederick  (A-F) 
model has been selected among other models to be the main model for the numerical simulations within 
the 
  project.  It  is  an  advanced  plasticity  model,  significantly  more  powerful  than  linear 
kinematic  models,  available  in  many  commercial  finite  element  codes  (e.g.  ABAQUS,  MARC  etc.). 
The  model  adopts  a  von  Mises  description  of  the  yield  surface,  and  assumes  a  nonlinear  kinematic 
hardening rule for the evolution of the back stress (i.e. the centre of the yield surface). The calibration 
procedure is based on material test results in Task 3.1, from coupon specimens for the material testing 
were extracted from the longitudinal and, mainly, from the hoop direction of elbow specimens.  The A-
F  model  is  calibrated  to  fit  the  material  cyclic  stress-strain  curves  and  the  appropriate  values  for  the 
corresponding  material  parameters  are  obtained,  considering  the  expected  strain  range  at  the  critical 
locations  of  the  specimens  under  consideration.  In  addition  to  the  A-F  model,  the  Tseng-Lee  model, 
based on the bounding surface concept is also used and calibrated from the coupon tests. Such a model 
is not incorporated in the commercial finite element codes (e.g. ABAQUS, MARC etc.) and, therefore, 
a special-purpose user subroutine has been developed. 

The  simulation  of  piping  component  experiments  conducted  in  WP3  (nozzles,  elbows,  flanges  and 
Tees), has been performed in Task 4.2, for the purses of developing and validating rigorous numerical 
tools.  To  achieve  a  good  level  of  accuracy,  nonlinear  finite  element  models  are  employed,  which 
employ  shell  elements  or  solid  elements,  and  the  nonlinear  constitutive  models  for  cyclic  plasticity, 
calibrated in Task 4.1. In particular, 

(a)  NOZZLES:  The  finite  element  models  were  developed  in  ABAQUS  using  8-node  reduced-
integration  solid  elements  (C3D8R),  considering  geometric  and  material  non-linearity.  Each  shell 
nozzle  reinforcement  type  was  investigated  under  longitudinal  and  transverse  loads,  as  in  tests.  By 
comparing the cyclic load-displacement curves as well as the points of failure in the test with those of 
the simulations, a good accordance was achieved. To estimate the number of cycles until failure, a quite 
simple approach was chosen, in terms of the equivalent plastic strains. 

(b)  FLANGED  (BOLTED)  JOINTS:  A  three-dimensional  model  of  the  flanged  joint  including  the 
gasket and the pipes using MARC. Both weld neck flange ring (ring and hub) and non-standard flanges, 
have  been  modelled  by  means  of 8-node  brick  element,  while the  gasket  sealing  element  employs  8-
node composite brick elements that allow adopting specific material model able to describe the gasket 
behaviour. The pipe has been modelled by means of shell elements, each bolt is modelled with a beam 
element,  the  head  of  the  bolt  is  modelled  as  rigid  and  pre-tension  is  taken  into  account.  The  spiral 
wound  gasket  has  also  been  modelled.  The  comparison  of  the  experimental  and  numerical  results 
demonstrated  a  good  correlation  in  terms  of  both  general  load-versus-stroke  evolution  and  specimen 
deflections. 

(c) ELBOWS: For the simulation of each conducted test, a numerical model has been developed, with 
the  exact  dimensional  characteristics  of  the  specimen  based  on  detailed  dimensional  measurements 
conducted prior to testing (thickness variability of the curved and the straight parts of the elbows has 
been taken into consideration). Weld over-thickness is also modeled. The finite element model has been 
developed  in  ABAQUS  with  eight-node  incompatible  mode  (C3D8I)  solid  elements.  The  material 
models  employed  are  the  Tseng-Lee  and  the  Armstrong-Frederick  models.  The  numerical  results  are 
reported  in  the  form  of  load-displacement  and  flattening-displacement  curves,  and  are  in  very  good 
agreement with the corresponding test data. Fatigue life has been estimated very successfully in terms 
of local strain variation at the crack location. 

(d) TEE-JUNCTIONS: Numerical models with solid elements in ABAQUS were used to simulate the 
tests  of  task  3.5.    The  comparison  of  the  predicted  load-displacement  curves  with  the  corresponding 
experimental curves shows that the adopted model is able to provide numerical results in good overall 
agreement with the experimental findings. Some differences can be attributed to minor slips at the end 
supports of the test specimens. These slips are more evident in the TJF tests due to their increased load 
resistance. In any case, the numerical models can predict the maximum and minimum resistance loads 
fairly well.  

11 
Finally,  in  Task  4.3,  the  above  numerical  tools  have  been  employed  in  performing  an  extensive 
parametric  numerical  study  of  the  components  under  consideration,  considering  a  wide  range  of 
parameters. The main issues related to this numerical investigation are described below: 

(a)  NOZZLES:  The  defined  parameters  are  (i)  plate  thicknesses,  (ii)  diameter  and  thickness  of  the 
attached nozzle and (iii) geometry of the reinforcements. Parametric studies have been performed for 
both  monotonic  and  cyclic  loading.  In  the  first  step,  monotonic  simulations  for  the  4  different  load 
directions for each type of reinforcement were conducted, which resulted in 240 monotonic simulations. 
The  influence  of  internal  hydrostatic  pressure  has  also  been  taken  into  account.  Subsequently,  a 
parametric numerical study has been conducted for cyclic loading in both longitudinal and transverse 
load  direction  and  each  type  of  reinforcement,  considering  an  increasing  amplitude  loading  (ECCS 
No.45  protocol).  The  results  have  been  presented  in  terms  of  relative  resistance  ratios  versus  partial 
ductility of the component. 

(b) FLANGED (BOLTED) JOINTS: The finite element models employed for the comparison with full 
scale  tests  of  the  flanged  joints  have  also  been  employed  to  perform  a  sensitivity  analysis  of  various 
load  scenarios  and  geometrical  parameters  on  flanged  joints.  Both  standard  and  non-standard  flanges 
have  been  analysed:  (a)  A  numerical  parametric  study  on  EN1092-1  type  11  welding  neck  standard 
flanged joints subjected to bending loading aimed at investigating the effect of the parameters such as 
internal pressure level, axial load and geometry on the behaviour of the bolted flanged connections. (b) 
In non-standard flanged connections, special attention was paid on the thickness of the flange, in terms 
of the relevant limit state, in both bending axial loading conditions. It has been demonstrated that non-
standard flanges can be used in several industrial piping applications for piping systems operating under 
both normal operating and seismic conditions. 

(c) ELBOWS: The behaviour of steel pipe elbows under strong cyclic in-plane loading has been studied 
numerically,  based  on  numerical  models  developed  to  simulate  the  experiments,  and  using  two 
appropriately calibrated plasticity models. Applying various ranges of end-displacement, strong cyclic 
in-plane  bending  loading  has  been  introduced  and  the  number  of  cycles  until  failure  has  been 
determined.  The  numerical  results  are  reported  in  the  form  of  load-displacement  and  flattening-
displacement curves. The effects of elbow geometrical characteristics on its mechanical behaviour have 
been studied. Furthermore, the case of out-of-plane bending has been examined; it was found that the 
formation of  local  buckling  results in  local strain concentration,  which  would  affect significantly 
the failure life of the elbow. 

(d)  TEE-JUNCTIONS:  Following  the  calibration  of  the  numerical  models,  an  extensive  parametric 
study  has  been  conducted.  For  the  Tee-junctions  under  consideration  the  following  parameters  of 
significant importance are taken into account: (a) tee-junction configurations, comparing those with a 
reinforcement plate with those that use a Tee-junction  fitting, (b) load direction, namely  in-plane and 
out-of-plane  bending  and  axial  loading  of the branch,  (c)  level  of internal  pressure, (d)  radius  of Tee 
fitting,  (e)  geometric  parameters,  such  as  diameter  and  thickness  of  branch  and  main  pipe,  (f) 
reinforcement  plate  geometry  (width  and  thickness  variations).  Both  configurations  are  capable  of 
enduring  high  seismic  loadings  with  the  fitting  being  the  strongest.  The  modes  of  failure  have  been 
identified and the capacity for the different geometries and loading directions determined. 

I.5.  WP5:  Seismic  behaviour  and  special  seismic  design  issues  in  liquid  storage 

tanks; development of relevant design guidelines/recommendations. 

The  main  objective  of  this  work  package  (WP5)  is  the  development  of  a  complete  set  of  design 
guidelines / recommendations for the seismic analysis and design of industrial liquid storage tanks. The 
guidelines are compatible with the existing Eurocode design framework, and extend the existing rules 
of Eurocode 8 (EN 1998-4).  

The work conducted in Task 5.1 consisted of examining several special issues on the seismic response 
of liquid storage tanks. The results of this investigation  can be considered as a background document 
for  the  seismic  design  guidelines  of  liquid  storage  tanks  to  be  presented  in  a  companion  report.  In 
particular, the following issues have been addressed in detail: (a) the number of sloshing modes to be 
considered  in  convective  action  and  their  combination  with  impulsive  action;  (b)  the  effects  of  shell 
deformation  on  the  seismic  response;  (c)  calculations  of  hoop  hydrodynamic  stresses;  (d)  uplifting 

12 
 
behaviour  of  unanchored  tanks;  (e)  buckling  at  the  top  of  the  tank;  (f)  elephant’s  foot  buckling  and 
behavior factor of liquid storage tanks; (g) nonlinear sloshing effects; (h) floating roof behaviour. 

Considering  typical  liquid  storage  tanks,  provided  by  our  industrial  partners,  the  investigation  of  the 
above issues was aimed at evaluating EN 1998-4 provisions (in comparison with those of API 650) for 
the  seismic  design  of  liquid  storage  tanks,  towards  proposing  possible  improvements/amendments  of 
EN 1998-4. The main conclusions from this investigation can be stated as follows: 

 

 

 

 

 

 

For the majority of tanks, consideration of only one sloshing (convective) mode is adequate 
for the calculation of the total seismic force. Furthermore, the combination of the convective 
force with the impulsive force should be conducted according to the SRSS rule. 
Using  a  systematic  two-step  methodology  for  analyzing  the  seismic  response  of  uplifting 
tanks  has  shown  that  the  EN  1998-4  provisions  provide  a  reasonable  yet  conservative 
approach for the increase of axial compression due to uplifting. 
A simple methodology is presented in order to design tanks against buckling at the top, due to 
liquid sloshing, using the concept of “equivalent” external pressure. 
The behavior (reduction) factor for the seismic design of liquid storage tanks is a controversial 
issue  that  has  not  been  examined  thoroughly  so  far.  Because  of  the  relatively  low  post-
buckling  strength  of the  shell tank,  the authors  would  be  quite  cautious in  using  a  behavior 
factor greater than 1.5 for the impulsive component of motion. 
There exists little information on current design practice for floating roofs; to avoid complete 
tank destruction, one should avoid seal damage to prevent fire ignition. 
Nonlinear sloshing effects on the seismic behavior have not been thoroughly investigated so 
far.  Limited  numerical  investigations,  reported  elsewhere,  indicated  that  their  effects  on  the 
total seismic  force  may  not  be  important,  but they  can  have  significant effects on  freeboard 
considerations. 

The design guidelines for the seismic design of liquid storage tanks have been developed in the course 
of Task 5.2, at an attempt to present a document that improves the existing provisions of EN 1998-4. 
The  structure  of  those  guidelines  is:  1  Introduction;  2  Relevant  Standards  and  Guidelines;  3  List  of 
Symbols; 4 Seismic Actions in Liquid Storage Tanks; 5 Seismic Resistance of Liquid Storage Tanks; 6 
Special seismic design provisions; ANNEX A: Hydrodynamic loading due to seismic action; ANNEX B: 
Unanchored tanks; ANNEX C: Shell Buckling Resistance. 

The  guidelines  can  be  employed  by  designers  of  liquid  storage  tanks  in  seismic  areas,  providing 
background information to existing  standards, and identification of ultimate limit states under seismic 
action, as well as amending and improving the existing provisions of EN 1998-4. 

I.6.  WP6:  Seismic  behaviour  and  design  of 

industrial  pressure  vessels; 

development of seismic design guidelines/recommendations. 

The  main  objective  of  this  workpackage  (WP6)  is  the  development  of  design  guidelines  / 
recommendations for the seismic analysis and design of industrial pressure vessels (vertical cylinders, 
horizontal cylinders and spheres). The guidelines concern various geometries and  are compatible with 
the existing Eurocode 8 (EN 1998) framework for earthquake-resistant design. 

The seismic behaviour of vertical and horizontal pressure vessels has been analyzed in Tasks 6.1, 6.2 
and  6.3,  using  pushover  and  incremental  dynamic  analysis  techniques.  The  failure  mechanisms  were 
discussed and possible failure criteria were proposed. The behavior factor q determined for the vertical 
vessels was found to be 1.5, lower than the value of 2 suggested by ASCE7 for such pressure vessel 
geometries. The effect of internal pressure was insignificant up to the point of maximum capacity of the 
vessel  but  beneficial  for  post-buckling  behavior.  For  the  horizontal  vessels  the  effect  of  reinforcing 
rings,  the  effects  of  length,  thickness  and  those  of  internal  pressure  were  studied  using  pushover 
analysis.  The  effect  of  internal  pressure  was  more  significant  when  loading  was  applied  in  the 
transverse rather than in the longitudinal direction due to the change of failure mode. Reinforcing rings 
are recommended for large pressure vessels or pressure vessels located in high seismicity areas. Caution 
shall be taken when deciding the area of the saddle reinforcement as the adjacent shell area is expected 
to fail in the case of no internal pressure. 

13 
 
Moreover, behaviour factors q for spherical pressure vessels have been determined and showed to be 
bigger than 2, evaluating conservatively the most critical earthquake from a set of 10 accelerograms. A 
q-factor  of  2.7  was  found  for  the  non-braced  system,  which  should  be  restricted  to  this  value  since 
dynamic  instability  phenomena  can  occurs  suddenly  using  higher  values.  For  the  braced  supporting 
structure a q-factor of 2.3 was identified, when limitation of the plastification to the bracings is desired. 
Otherwise a higher value of 3 has been suggested. 

The guidelines for seismic design of pressure vessels are developed in Task 6.4 considering cylindrical 
pressure  vessels  that  are  vertical  and  horizontal  or  of  spherical  shape.  They  consider  ground  level 
supports,  elevated  supports  and  supports  by  large  structures.  Recommendations  for  achieving  ductile 
response  are  given  by  means  of  global  design  concepts  and  local  detailing  rules.  The  guidelines  for 
seismic design are structured as follows: Introduction; Relevant standards; List of symbols; Definition 
of limit states; Seismic Input; Seismic analysis; Seismic design; Design Examples; ANNEX A: Supports 
by large structures. 

The  developed  guidelines  are  aimed  for  use  by  pressure  vessel  designers.  They  are  meant  to  be 
considered  in  addition  to  existing  seismic  design  rules  for  steel  structures  which  -  to  large  extend  - 
remain  valid  for  pressure  vessels  and  in  particular  for  their  supporting  structures.  The  guidelines  are 
also accompanied by specific design examples. 

I.7.  WP7:  Seismic  behaviour  and  design  of  industrial  piping  systems  and 

components; development of seismic design guidelines/recommendations. 

The main objective of WP7 is the development of design guidelines for the seismic analysis and design 
of industrial piping, including their support systems. Special emphasis has been given to the design of 
an adequately dissipative support system, to reduce the seismic action on the piping. The guidelines are 
compatible with the existing Eurocode 8 (EN 1998) framework for earthquake-resistant design. 

Extensive work has been conducted in task 7.1, to develop background documents for the guidelines. 
The study aimed at identifying all the design steps of piping systems and all design aspects that need to 
be clarified. This was made through a representative case study, analysed according to both European 
(EN13480:3) and  American standards (ASME B31.3) devoted to piping systems.  With reference to a 
typical piping system the comparison between EN13480 and ASME B31.3 standards showed that the 
two codes provide similar indications for the evaluation of the seismic response of pipelines supported 
by a pipe-rack. The usual way of designing piping systems is based on the allowable stress approach, so 
that  the  structure  is  considered  elastic.  Consequently  only  an  operating  basis  earthquake  condition 
(OBE) can be taken into account for design. Conversely, the modern approach to the seismic design of 
structures  is  to  differentiate  serviceability  from  ultimate  limit  states.  This  latter  condition  could  be 
represented  by  the  Safe  Shutdown  Earthquake (SEE).  It  was  identified  that a  proper definition of the 
deformation  capability  of  the  pipes  and  fittings  (ovalization,  etc.)  in  the  plastic  range  would  be 
necessary, and on this aspect the research has not reached solid conclusions. However, the advantages 
would  be  several  and  first  of  all  the  design  optimisation.  For  the  case  study  under  consideration,  the 
acting  moment  is  well  below  from  the  moment  capacity  of  the  bolted  joint  of  the  piping.  This 
conservatism seems to be in contrast with the modern performance based-design approach, for which a 
certain level of yielding in the system is allowed according to a specific performance.  

The  guidelines have been developed in task 7.2, and consist  of  a  concise  document compatible  with 
the  existing  Eurocodes  design  framework,   which  includes  some  useful  indications  on  seismic 
analysis  of  piping  systems.  The  document  follows  the  basic  principles  of  earthquake  engineering 
expressed  in  EN1998:1, extending  the  existing  rules  to  industrial  piping  components  and  systems. 
The  guidelines  are  organized  in  the  following  chapters:  Scope;  Definitions;  Reference  documents, 
Symbols  and  Abbreviations,  Methods  and  relevant  parameters  for  seismic  analyses;  Seismic  input; 
Modeling  of  piping  systems;  Dynamic  seismic  analyses;  Limit  states  (including  results  obtained  from 
the testing programs); Recommendations. 

The  guidelines  will  be  used  by  designers  of  industrial  piping  systems,  providing  background 
information to the existing standards, description of critical steps and parameters in seismic design of 
piping systems, recommendations for modelling of piping systems and the interaction with supporting 
substructure and identification of ultimate limit states under seismic action.  

14 
 
I.8.  WP8: Coordination of the project. 

The  main  objectives  of  WP8  have  been  the  establishment  of  an  efficient  management  process;  the 
definition of standard protocols; the performance of SWOT analysis; and the continuous monitoring of 
the project. All coordination meetings were regularly carried out, the FTP server directory was created 
by  UThessaly  and  continuous  exchanges  of  information  favoured  objectives  and  relevant  project 
outcomes.  Regarding  testing  of  components  in  WP3,  common  standard  testing  protocols  have  been 
decided  for  assuring  quality  of  test  results.  In  addition,  the  range  of  geometric,  material  and  loading 
parameters  to  be  used  in  the  numerical  studies  in  WP4  have  been  determined.  Furthermore,  the 
evaluation  of  the  project  in  terms  of  Strength,  Weakness,  Opportunity  and  Threat  allowed  a  global 
evaluation of the research project both in terms of activity coordination and work effectiveness of each 
research unit. Intermediate annual reports and deliverables enabled internal and external monitoring of 
the project advances, and of the degree of fulfilment of objectives set for various phases.  

I.9.  Conclusions and main results 

INDUSE

The 
  project  has  been  a  pioneering  project  within  the  RFCS  research  program.  It  combined 
experimental, numerical and analytical tools to conduct high-level research in the inter-disciplinary area 
of  seismic  safety  of  industrial  facilities.  Extensive  experimental  and  numerical  work  has  been 
conducted, to investigate the mechanical behavior of critical components; their behaviour under strong 
cyclic loading has been shown well above the requirements of existing design provisions. In addition, 
 project resulted in the development of design guidelines/recommendations (a) expanding 
the 
EN 1998-4 provisions towards an integrated seismic design of liquid storage tanks and (b) extending the 
applicability  of  Eurocode  8  seismic  design  concepts  for  the  cases  of  industrial  pressure  vessels  and 
attached equipment and piping systems including their support systems. The results and deliverables of 
INDUSE

 are both novel and unique, resulting in  

INDUSE

 

 
 

better understanding of structural behaviour of tanks, pressure vessels and piping under strong 
cyclic (seismic) loading,  
more reliable definition of ultimate limit states for critical industrial components  
significant improvement of seismic design state-of-the-art 

and offering the necessary scientific background for Code Drafting Committees for improving existing 
design standards, towards safer design of industrial power plants and chemical/petrochemical facilities. 

15 
 
 
 
II. Scientific description of the results 

Objectives of the project 

 

 

 

 

 

Critical evaluation and comparison between currently available seismic design provisions in 
European and American standards (WP1), considering specific case studies.  
Determination  of  seismic  actions  on  industrial  systems  (WP2),  taking  into  account  the 
dynamic behaviour of each structural system. 
Testing of critical piping components (nozzles, flanged connections, Tees and elbows) under 
strong cyclic loading (WP3). Pseudo-dynamic testing of a piping system. 
Performance of extensive numerical studies on critical piping components, based on rigorous 
FE simulations, covering a wide range of geometric, material and loading parameters (WP4). 
Development  of  design  guidelines/recommendations,  expanding  EN  1998-4  provisions 
towards  an  integrated  seismic  design  of  liquid  storage  tanks  (WP5)  and  extending  the 
applicability  of  Eurocode  8  (EN  1998)  concepts  for  the  cases  of  industrial  pressure  vessels 
and attached equipment (WP6), and piping systems including their support systems (WP7). 

Comparison of initially planned activities and work accomplished 

All  the  above  objectives,  stated  in  the  Technical  Annex,  have  been  successfully  accomplished  in  the 
course of the relevant work packages of 

INDUSE

 project. 

II.1.  Work  package  1  –  Comparison  of  European  and  American  design  standards 
for  the  seismic  design  of  industrial  liquid  storage  tanks,  pressure  vessels  and 
piping and development of available test data base 

The objectives of WP1 are summarized as follows: 

 

 

 

Collection and evaluation of cyclic loading (static and dynamic) test data on critical industrial 
components of industrial systems available in the literature.  
Collection  and  evaluation  of  damages  in  industrial  equipment  structures  subjected  to  strong 
earthquake events 
Comparison  of  currently  used  European (EN  and  national) standards  and  American  (API  or 
ASME) standards, in typical industrial equipment structures, namely (a) liquid storage tanks, 
(b)  pressure  vessels  and (c)  industrial  piping,  to  identify  main  differences  in  seismic  design 
procedures.  

II.1.1. 

Task 1.1: Collection and evaluation of available relevant experimental data 

II.1.1.1.

Experiments on piping components and piping systems 

In this critical evaluation of data, experiments on piping components and piping systems are described 
and discussed. The following experimental testing programs on piping components and piping systems 
are described, evaluated and extensively discussed. 

1. 

2. 

3. 

4. 

5. 

Berkeley National Laboratory tests, UK,  

University of Liverpool tests I 

University of Liverpool tests II 

EPRI-NRC Piping and Fitting Dynamic Reliability Program 

ETEC System Tests 

6.  Westinghouse Hanford Test 

7. 

8. 

EPRI Prototype Tests 

Japanese High Level Tests 

17 
 
 
 
 
9. 

University of Akron Studies 

10. 

Japanese Vibration Tests 

II.1.1.2.

Experiments on liquid storage tanks 

The dynamic behavior of ground-based liquid storage tanks had been the subject of extensive studies at 
University  of  California  Berkeley  during  70’s-80’s.  Theoretical  and  experimental  investigations  had 
been conducted to seek possible improvements in the  design of such tanks to resist earthquakes [46], 
[15],  [37],  [36].  These  experiments  were  mainly  directed  towards  studying  sloshing  behavior  and 
buckling of the shells.  

More recently, in the same university, a well-organised series of tests took place [48]. Experimental and 
analytical studies were conducted on seismically isolated and fixed-base liquid storage tanks. Shaking 
table tests were performed on vertical cylindrical open-top steel tanks [48]. Extensive parameter studies 
were performed, including comparisons of fixed-base and isolated model response to simulated ground 
motions of varying magnitude, time scale, and combination of components.  

Finally,  in  [43]  the  effectiveness  of  the  base  isolation  on  steel  storage  tanks  has  been  investigated 
through numerical models and then checked by shaking table tests on a reduced scale model of a real 
steel tank,  typically  used  in  petrochemical  plants.  In  the  experimental  campaign  the floating  roof  has 
also been taken into account. 

II.1.2. 

Task  1.2:  Collection  and  evaluation  of  reported  earthquake  damages  in 
industrial facilities 

The examination of the seismic impact on industrial facility systems, and it is extremely valuable as one 
attempt to assess the probable risks in a future earthquake. The following review is in the framework of 
Task 1.2, and can serve as a guide to identify potential problems and weaknesses. In this report a few 
notable examples of seismic damages are given where the impact was more severe. 

The collection and evaluation of typical damages in industrial equipment structures subjected to strong 
earthquake events has been performed in task 1.2, for the following earthquake events: 

1. 
2. 
3. 
4. 
5. 
6. 

Imperial Valley Earthquake, California, USA, October 15, 1979 
Coalinga Earthquake, California, USA, May 2, 1983 
Northridge Earthquake, Los Angeles, USA, January 17, 1994 
Kocaeli Earthquake, Turkey, August 17, 1999 
Chile Earthquake, March 3, 1985 
Kobe Earthquake, January 17, 1995 

The main conclusions from this investigation can be briefly summarized as follows: 

 

 

 

 

 

 

 

Earthquakes may result in very significant damages to the industrial facilities. 

Main damages occurred to liquid storage tanks. 

The principal tank damage is elephant’s foot buckling at the tank bottom. 

In addition, damages to the top of the tanks may occur due to sloshing effects. 

Piping equipment may also exhibit significant damages, due to inadequate flexibility. 

In many instances, fire that occurred (especially in tanks) resulted in a complete destruction of 
the tank. 

Failure of pipe support equipment has also occurred during the earthquake. 

The above evaluation is taken into account in the next stages of the project 

18 
 
 
 
II.1.3. 

Task 1.3: Basic seismic structural design of typical liquid storage tanks 

The present task focuses on the seismic design of liquid storage tanks. Five case studies are considered 
and designed against seismic loads according to the following international design standards for liquid 
storage tanks.  

 
 
 
 
 

Case Study #1: 27.4-m-diameter liquid storage tank  
Case Study #2a: 18-m-diameter liquid storage tank  
Case Study #2b: 6-m-diameter liquid storage tank  
Case Study #2c: liquid storage tank with floating roof  
Case Study #3: broad liquid storage tank  

Their seismic design has been performed using European and American standards. Seismic actions have 
been calculated, and the necessary verification checks were conducted. Inconsistencies and issues that 
require  further  consideration  have  also  been  identified.  The  conclusions  from  this  study  are  used  in 
WP2 and WP5 and can be summarized as follows:  

1. 
2. 
3. 

4. 

5. 

6. 

7. 
8. 

9. 

There is a significant revision of API 650 11th edition with respect the 10th edition. 
EN 14015 in seismic design follows the old edition of API 650, needs also revision. 
Both specifications API 650 (2007) and EN 1998-4 follow closely the hydrodynamic solution 
(convective/impulsive) for the contained liquid. 
API 650 is a more design-oriented code, whereas EN 1998-4 is a document that contains very 
good information, but sometimes difficult to apply. 
Section A.3 on deformable tanks needs significant revision. The extensive rules stated in this 
section are almost impossible to apply for design purposes. 
There  are  significant  differences  in  the  overturning  moment  (ring-wall  and  foundation)  and 
the base shear, because of the significant difference in the reduction (behavior) factor on the 
seismic design acceleration. 
There is a difference in the formula for sloshing wave height (freeboard). 
API 650 has explicit formulae for hoop stresses whereas EN 1998-4 does not. The accuracy of 
those formulae needs to be investigated. 
API  650  adopts  a  SRSS  impulsive/convective  combination,  whereas  EN  1998-4  adopts 
summation of design values. 

10.  EN  1998-4  needs  some  revisions  in  section  A.9  (uplifting  of  unanchored  tanks)  to  reach  a 

more user-friendly form. 

11.  Direct  comparison  between  EN  1998-4  section  A.9  and  API  650  E.6.2.1  for  the  effects  of 

uplifting is not possible; a more in-depth investigation is necessary. 

II.1.4. 

Task  1.4:  Basic  seismic  structural  design  of  typical  cylindrical  and  spherical 
pressure vessels 

In  this  Task,  a  comparison  of  currently  used  European  and  American  standards  for  pressure  vessels 
design  is  being  carried  out.  In  particular,  four  case  studies  are  considered,  which  constitute  typical 
pressure vessels of petrochemical facilities. 

  Case Study #4: spherical pressure vessel  
  Case Study #5: vertical pressure vessel on a skirt 
  Case Study #6: horizontal pressure vessel on saddle supports 
  Case Study #7: vertical pressure vessel on a steel structure  

The  analysis  and  design  of  the  vertical  pressure  vessel  supported  on  a  skirt  and  horizontal  pressure 
vessel supported on saddles indicated the following issues: 

1.  Both  specifications  EN  13445  and  ASME  VIII  do  not  contain  specific  provisions  for  the  seismic 
analysis  of  pressure  vessels  and  the  determination  of  seismic  forces.  Such  guidance  should  be 
developed and incorporated in the proposed guidelines. 

2.  Both specifications EN 13445 and ASME VIII indicate an increase of stress allowables for seismic 
loading. In EN13445 the safety factor can be increased up to the one used for the testing load cases 
for both tension and compression. In ASME VIII Div.1 a 20% increase is allowed in tension, while 

19 
 
in ASME Div.2 no such increase is specified apart from the compression allowable used in Zick’s 
analysis. 

3.  The  main  conclusion  from  the  above  design  and  analysis  of  the  two  pressure  vessels  is  a  lack  of 
. The only source for choosing a 
guidance for choosing an appropriate value of the behavior factor 
value for the behavior factor for pressure vessels is Table 15-4.2 of ASCE 7-05, but these values are 
not  justified.  More  research  work  is  necessary  towards  determining  appropriate  values  of  the 
behavior factor 

. 

4.  The calculation of the natural period of the vessel depends on the flexibility of the supports, and may 
affect significantly the seismic force. This causes some uncertainty in calculating the natural period 
especially in the horizontal vessel on saddles supports. Therefore, it would be helpful to specify one 
method to calculate the natural period to be used for design. 

5.  The design procedure in ASME VIII is based on local stresses calculated from Zick’s analysis. On 

the other hand, the design procedure in EN 13445 is based on stress resultants. 

The  analysis  and  design  of  the  vertical  pressure  vessel  on  a  steel  structure  leaded  to  the  following 
conclusions. 

1.  First  observation  is  that  the  analysis  following  EN1998-1  Clause  4.3.5  for  appendages  arrives  to 

more conservative Total Base Shear Values.  

2.  On the contrary analysis following EN1998-1 Clause 3.2.2 for structures supported on ground level, 
lead to lower Total Base Shear Values. This result seam logic and it is something that was expected 
to arrive to higher Base Shear value on a certain height from the ground due to higher acceleration 
value. 

3.  Comparing the results from the above two with the results from integrated method (Col. 2), we see 
that  Base  Shear  value  is  between  the  above  two  cases  (Col.  4  &5).  We  are  not  in  position  to  say 
which values are more realistic even though with Integrated Methods Analysis we could have more 
realistic  parameters  concerning  structure  general  geometry,  stiffness  that  may  affect  on  Floor 
Acceleration. The matter should be further investigated. 

4.  Additionally, are the results following ASCE7-05 in between the results for Total Base Shear from 
the  above  two  methods.  That  method  does  not  take  into  consideration  the  interaction  between 
pressure vessel and structure natural frequency.  

5.  Concerning  the  significant  difference  in  shear  that  appears  in  –y  direction  at  the  analysis  using 
Integrated  Method  we  assume  that  phenomenon  could  be  explained  because  building  overall 
geometry (B.R.F. on –y direction), or perhaps local eccentricities of the vessel with respect of on –x 
and –y axes. 

II.1.5. 

Task  1.5:  Basic  seismic  structural  design  of  typical  piping  system  and  its 
support structure 

In this Task, a comparison of currently used European and American standards for piping system design 
is being carried out. In particular, two case studies are considered. 

  Case Study #8: piping system on a pipe-rack  
  Case Study #9: piping system on a steel support structure  

First,  a  piping  system  on  a  pipe-rack  is  considered  and  designed  against  seismic  loads  according  to 
European  [25]  and  American  standards  (ASME  B31.3)  for  piping  systems.  The  comparison  between 
EN13480 and ASME B31.3 standards has shown that the two codes provide similar indications for the 
evaluation  of  the  seismic  response  of  pipelines  supported  by  a  pipe-rack.  From  the  analysis  of  the 
results  the  following  conclusions,  considered  valid  for  both  the  standards,  can  be  drawn.  Moreover, 
some recommendations are provided. 

1.  Dynamic  coupling.  ASCE-07  [5]  provides  a  simple  rule  to  establish  when  the  dynamic  coupling 
between  pipes  and  supporting  structure  would  be  considered.  It  is  based  only  on  a  weight  ratio. 
Actually, other ingredients should be included; for example the vibration periods range of the pipe. 
In the analyzed case the single pipe have a limited weight and therefore short vibration periods. This 
means that, as already shown, the dynamic coupling between the pipes and the pipe-rack is limited, 

qq20 
 
whereas  the  relative  displacements  between  the  pipe  supports  would  represent  a  more  relevant 
effect. The pipes provide a sort of static coupling between the transverse frames. In fact, the models 
with  and  without  the  pipe  stiffness  contribution  have  more  or  less  the  same  period,  whereas  the 
excited  mass  is  differently  distributed  between  the  several  vibration  modes.  Therefore,  the 
consequence of considering the structural contribution of the pipes is the introduction of a sort of a 
rigid floor effect that couples the horizontal movements of the nodes of the same floor. 

2.  In-Structure spectra. The in-structure (floor) spectra suggested by prEN1998:1 and ASCE-07 differs 
for a term that depends on the pipe-structure  frequency ratio, included only in the European code. 
This aspect is strictly related to the previous observation on the dynamic coupling. The exclusion of 
the dynamic interaction could introduce a high error in the evaluation of the amplification factor of 
in-structure  spectra.  Because  usually  the  frequency  of  the  piping  system  is  relatively  high  with 
respect to that of the supporting structure, the amplification effect would be relatively low. In any 
case,  to  cover  all  the  possible  cases,  the  use  of  the  European  formula  for  in-structure  spectra  is 
highly recommended.   

3.  Modeling  of  the  pipes.  In  the  analysis  both  beam  element  and  shell  elements  have  been  used  to 
model elbows. The comparison has shown the reliability of the more simple elements based on the 
beam theory, at least for the simple case of elbows. For more complicated situations the use of finite 
shell elements should be considered.  

4.  Behavior  factor.  The  behavior  factor  q  is  usually  indicated  by  the  codes.  In  particular  the  values 
provided  for  piping  systems  by  ASCE-07  seems  to  be  too  much  high  (R=612).  This  in  contrast 
with  the  idea  of  avoiding  yielding  phenomena  in  the  pipes,  in  order  to  maintain  operational 
conditions  also  in  case  of  strong  seismic  events.  Certainly,  the  contribution  of  the  supports  in 
providing dissipation capability could be taken into account, but on the basis of several investigation 
found in the literature this exclude the possibility of having so high values of R.  

5.  Performance based-design of pipes. The usual way to design piping systems is based on allowable 
stress approach. This means that the structure is considered elastic and the seismic action in reduce 
of a safety coefficient equal to 1.4. Consequently only operating basis earthquake condition (OBE) 
can be taken into account for the design. On the contrary, the modern approach to the seismic design 
of  structure  is  to  differentiate  serviceability  from  ultimate  limit  states.  This  latter  could  be 
represented for example by the Safe Shutdown Earthquake (SEE).  

In case study no 9, two different piping and support structures has been selected from a project, situated 
in  an  area  of  moderate  seismic  activity  (Algeria),  which  was  completed  by  Tebodin  in  2009.  Both 
simplified static equivalent (‘uniform load method’) and dynamic calculations were made in accordance 
with American (ASCE7) and European (EN 1998-1/3 and EN 13480) earthquake design standards to 
identify differences in approach, design seismic response spectra and results. The analysis of the piping 
systems leaded to the following conclusions: 

1.  The design seismic spectra according to EN1998-1 and ASCE-are similar and therefore yield similar 

stresses, forces and bending moments. The peaks occur roughly in the same periodicity. 
2.  The influence of the modeled steel support structure on the calculation results was limited. 
3.  For typical plant piping systems, having periods up to 0.3 sec., the uniform load method yield more 

conservative results than the dynamic seismic calculations.  

4.  Long above ground pipeline systems on many supports may however have many modes of vibration 
including lower frequencies and the a dynamic analysis may reveal higher stresses then calculated 
with the uniform load method. 

5.  The existing earthquake standards are mainly different with respect to the design response seismic 
spectra.  Design  and  correct  modeling  however  could  prove  to  be  of  greater  importance  on  the 
dynamic  calculation  results  then  the  difference  in  response  spectra.  It  is  recommended  to  include 
guidelines for design and modeling in the next revisions of these standards. 

21 
 
II.2.  Work package 2 – Determination of seismic forces on industrial structures and 

critical components and preparation of experiments 

The objectives of this workpackage (WP2) have been: (a) the development of systematic procedures to 
calculate  seismic  action  on  tanks,  pressure  vessels  and  piping,  using  advanced  analysis  methods, 
considering the particularities of the above systems; (b) investigation of the effects of nonlinear support 
system response on the seismic behaviour of pressure vessels and piping; examination of the influence 
of  dissipative  support  systems  towards  seismic  force  reduction;  (c)  Determination  of  cyclic  loading 
earthquake action on selected critical components, considering strong seismic input, and employing the 
above  analysis  procedures;  (d)  Specification  of  specimen  geometry  to  be  tested  in  the  subsequent 
experimental program together with the detailed specification of base material and welding procedures. 

II.2.1. 

Task 2.1: Seismic forces in vertical cylindrical liquid storage tanks 

The work in this task examines some special issues on the structural behaviour of upright-cylindrical 
liquid storage tanks, with emphasis on seismic loading issues. The tanks under consideration are widely 
used in industrial facilities and for water storage. Two main design standards are considered: EN 1998-
4 and Appendix E of API 650. There are significant differences between the two specifications, which 
are due to the fact that there exist several controversial issues on this subject, open to further research. 
These issues are (a) the number of modes necessary to estimate accurately the convective seismic force 
due to the hydrodynamic behaviour of the liquid containment; (b) the appropriate combination of the 
impulsive  and  the  convective  component  of  seismic  force;  (c)  the  uplifting  behaviour  of  unanchored 
tanks, with emphasis on the base plate behaviour and the increase of meridional compression; (d) the 
choice  of  an  appropriate  reduction  (behaviour)  factor  for  calculating  both  the  impulsive  and  the 
convective force; (e) the calculation of hydrodynamic hoop stresses due to liquid hydrodynamic motion; 
(f) the design of tanks against buckling at the top due to liquid sloshing; (g) the importance of nonlinear 
wave sloshing effects. 

In this study, the main conclusions can be stated as follows: 

  For the majority of tanks, consideration of only one sloshing (convective) mode is adequate for the 

calculation of the total seismic force.  

  The combination of the convective force with the impulsive force should be conducted according to 

the SRSS rule, which in accordance with the design provisions of API 650, Annex E. 

  Numerical results show that the EN 1998-4 provisions provide a reasonable yet conservative of the 

increase of axial compression due to uplifting. 

  A simple methodology is presented in order to design tanks against buckling at the top, due to liquid 

sloshing, using the concept of “equivalent” external pressure. 

  The behavior (reduction) factor for the seismic design of liquid storage tanks is a controversial issue 
that has not been examined thoroughly so far. The striking difference between the values of API 650 
and EN 1998-4 is notable. Because of the unstable post-buckling behavior of thin-walled shells, the 
authors  would  be  quite  cautious  in  using  a  behavior  factor  greater  than  1.5  for  the  impulsive 
component of motion. 

  Nonlinear  sloshing  effects  on  the  seismic  behavior  have  not  been  thoroughly  investigated  so  far. 
Limited numerical investigations, reported elsewhere, indicated that their effects on the total seismic 
force may not be important, but they may affect freeboard considerations. 

II.2.2. 

Task 2.2: Seismic forces in pressure vessels 

This task  focuses  on  the  seismic  analysis and  design  of  horizontal-cylindrical  and  vertical-cylindrical 
pressure vessels. First, a comparison of the seismic provisions in ASME B&PV code and in EN 13445-
3 is conducted for two specific case studies. The two case studies are considered in the present task; (a) 
a vertical-cylindrical slender pressure vessel on skirt and (b) a horizontal-cylindrical pressure vessel on 
saddles. Both pressure vessels are typical vessels for petrochemical facilities. In addition, a numerical 
investigation  is  conducted,  aimed  at  proposing  a  reliable  value  of  the  behavior  factor  for  those 

22 
 
 
structures. Those issues are examined numerically, simulating the above pressure vessels with nonlinear 
finite  element  models.  Finally,  a  methodology  for  determining  seismic  forces  due  to  liquid  sloshing 
horizontal-cylndrical  vessels is outlined. The results of  the present  study  can  be  used towards  a  safer 
and more reliable design of industrial pressure vessels under seismic action. 

Two case studies are considered in the present task; (a) a vertical-cylindrical slender pressure vessel on 
skirt  and  (b)  a  horizontal-cylindrical  pressure  vessel  on  saddles.  Both  pressure  vessels  are  typical 
vessels for petrochemical facilities. 

The  seismic  analysis  and  design  of  these  pressure  vessels  indicated  that  more  detailed  information 
regarding  the  estimation  of  the  behavior  factor  for  these  structures  should  be  included  in  the  design 
codes,  as  well  as  more  detailed  guidelines  for  the  determination  of  the  seismic  forces.  An  attempt  to 
estimate the behavior factor for these structures has been conducted using advanced numerical tools, for 
pushover analysis and incremental dynamic analysis. The pushover analysis indicated that the capacity 
of the presure vessels is quite higher than the required strength in seismic loading. On the other hand, 
using  incremental  dynamic  analysis  for  the  vertical-cylindrical  pressure  vessel,  the  value  of  behavior 
factor calculated is somewhat below the corresponding value suggested by American standard ASCE 7. 

II.2.3. 

Task 2.3: Seismic forces in industrial piping systems 

In this Task, the seismic analysis of a piping system has been studied giving particular emphasis on the 
calculation of the seismic action.  

To analyze the main differences of the possible seismic action models, an application to the case study 
8,  widely  studied  using  the  current  codes  for  seismic  analysis  of  piping  systems,  has  been  used.  In 
particular, both static and dynamic analyses have been carried out, using the response spectrum method 
and time-history analysis. From the results the following main conclusions can be drawn: 

  The problem of dynamic interaction between pipes and pipe-rack could play an important role in the 
evaluation of the seismic analysis of a piping system. The classical primary-secondary systems may 
be applied with difficulty to a piping system because they are usually based on single multi-support 
secondary  systems,  whereas  a  piping  represents  a  multi  multi-support  secondary  systems.  In 
literature few works have been published in the past on this subject and more effort should be spent.  

  The  in-structure  spectra  suggested  by  the  code  are  used  as  mean  to  evaluate  the  maximum 
acceleration to apply to pipes for evaluating the maximum stress. This approach in most of cases is 
certainly on the safe side, because overestimates the stress level of the pipes. But in case of complex 
structural situation in which a piping system may be realized, this could not be strictly true. To avoid 
problem, the generation of floor-spectra is recommended. This approach, often followed in design of 
nuclear plant equipments, consists in performing a time-history analysis and evaluating the absolute 
acceleration at each floor and the floor-spectrum is carried out. Thus, it is a direct approach difficult 
to be generalized. This approach has also been followed for the case study #8 using a set of three 
accelerograms,  which  has  highlighted  the  important  role  of  the  support  structure  in  filtering  the 
seismic  action.  Therefore,  further  analysis  will  be  dedicated  to  the  study  of  formulas  for  the 
evaluation of floor-spectra.  

  The analyzed case study #8 has evidenced that in some cases the problem of a correct evaluation of 
the piping stress level is not univocally related to a correct determination of the acceleration at floor 
where the pipes are placed. In fact, the relative displacements between supports of the same structure 
could play an important role on the stress level. This means that for a correct evaluation of the stress 
level  in  the  pipes  a  multi-support  excitation  should  be  used,  instead  of  the  classical  uniform 
excitation. To clarify this aspect further analysis will be carried out.  

  Another problem directly related to the previous ones is the determination of the response factor q 
either if we use response spectra analysis or floor-spectra analysis. The results shown in literature 
indicate that the q factor imposed by the American code is highly overestimated and that the value 
suggested  by  other  standards  and  guidelines  as  American  Lifelines  Alliance  proposal  seem  to  be 
more reasonable: the value of q is comprised within the range 1.-3.5. Therefore, further analysis will 
be  dedicated  to  this  particular  aspect  that  has  a  strong  influence  on  the  seismic  design  of  the 
structures. 

23 
 
II.2.4. 

Task 2.4: Dissipative support systems 

The investigation of nonlinear support system response on the seismic behaviour of pressure vessels, for 
the  purpose  of  reducing  of  the  seismic  force,  has  been  investigated  in  this  task,  using  a  simplified 
model. The investigations performed within the study focus on: verification of the dependency of the 
floor  response  spectra  on  the  fundamental  period  of  the  supporting  structure  and  dependency  on  the 
nonlinear (dissipative) behaviour of the supporting structure. 

The  first  stage  of  the  verification  of  the  floor  response  spectra  provided  by  the  codes  comprises  the 
calculation  of  the  responses  of  ideal  SDOF-models  representing  structures  supporting  industrial 
components. Additional results focused on the dependency of the extreme response spectra values on 
the  amplification  factor  of  the  acceleration  and  on  the  maximum  ductility  ratio  achieved  during  the 
response.  For  each  frequency  two  examples,  representing  the  most  different  responses  for  different 
accelerations, were obtained.  

The results obtained by the simulations generally confirm the assumptions that the fundamental period 
(resp.  corresponding  ground  acceleration  spectrum)  and  the  dissipative  behaviour  of  the  supporting 
structures  do  have  a  significant  influence  on  the  floor  response  spectra.  Dissipative  behaviour  of  the 
supporting structure can reduce the floor spectra significantly; also the response spectra of long period 
structures  with  low  spectral  ground  acceleration  are  lower  than  those  of  short  period  structures. 
Compared to the provisions given by EC 8, the determined spectra of short period supporting structures 
are  partially  significantly  larger.  On  the  other  hand  the  provisions  of  EC  8  for  long-period  structures 
achieving a moderate ductility ratio are significantly higher than the expected floor response spectra. In 
this context however it needs to be noted, that in order to benefit from the dissipative behaviour of the 
supporting structure this structure must undergo plastification, i.e. possible overstrength of the material 
may play a significant role. 

In conclusion it can be stated that the floor response spectra shall undergo a systematic revision in order 
to enhance both – the safety and the effectiveness of the code provisions. A proposal for a new function 
for the floor response spectra cannot be provided on the basis of the results presented above only. For 
the  time  being  the  only  possibility  for  the  achievement  of  reasonable  results  is  the  modelling  of  the 
supporting structure together with the components. Such a possibility is explicitly included in the ASCE 
7-05. 

II.2.5. 

Task 2.5 Determination of loading amplitudes on critical components 

The present task is aimed at using advanced numerical tools in terms of nonlinear finite elements for 
determining loading on critical piping components (nozzles, T-junctions, elbows, flanged connections) 
in  a  piping  system,  in  the  course  of  strong  seismic  input.  Towards  this  purpose,  a  typical  industrial 
piping system is considered, which comprises all the above components (nozzles, T-junctions, elbows, 
flanged connections), to be used later in WP3 for experimental testing. The main purpose of developing 
this  model  is  the  correlation  of  the  loading  amplitudes  applied  in  WP3  on  the  piping  component 
specimens, with the level of seismic input applied on this typical system. 

II.2.6. 

Task  2.6  Specification  of  procedures  for  specimen  manufacturing  and  base 
material supply 

For uniformity among the various experiments of the partners, it has been decided to use P355N quality 
material for pressure pipes and fittings according to EN10216.03 or equivalent. For the plates P355N 
quality  according  to  EN10028.03  was  selected.  The  selection  of  this  material  against  other  materials 
with lower grade (e.g. grade 235 or grade 265) used in those applications, is based on the fact that the 
results of the present project will be mainly applicable for new structures to be conducted in the next 
years, and therefore, the use of an “obsolete” material may not be appropriate. 

The  material  has  been  purchased  and  manufacturing of the test specimens  for WP3  has  started.  Each 
partner  was  responsible  for  the  supply  of  the  base  material  and  manufacturing  of  the  specimens. 
However,  for  achieving  uniformity  between  the  specimens,  it  has  been  agreed  that,  UThessaly  and 
EBETAM have a common supplier and manufacturer for the Tee-joint piping connections. In addition, 
TUDelft  and  CSM  will  have  a  common  supplier  of  the  elbows.  An  extra  Tee  joint  was  ordered  by 

24 
 
 
UThessaly/EBETAM  and sent to  UTrento  for  the  pseudo-dynamic  testing.  Furthermore,  a  number  of 
elbows  from  the  TUDelft/CSM  provider  and  manufacturer  are  also  sent  to  UTrento  for  the  same 
purpose (pseudo-dynamic testing). The straight pipe elbows have been provided by the corresponding 
provider as well, for the sake of uniformity. 

Uniform welding procedures have been agreed where possible, similar to common industrial practices. 
EBETAM  was  responsible  to  write  the  welding  procedure  specification  for  the  various  geometries. 
Welding  of  the  non-standard  flanges  manufactured  by  the  UTrento  in  WP3  will  be  detailed  in  the 
corresponding chapter. 

Welding  procedure  will  be  certified  according  to  EN  15614.01  and  welder’s  performance  will  be 
certified according to EN287.01 for manufacturers of the test specimens. All samples will be inspected 
for surface defects to ensure the quality of the welded samples. Visual Examination will be performed 
according to EN 970. Magnetic Particle Testing (EN 1290) or Liquid Penetrant Testing (EN 571-1) will 
also be performed and a test report will be submitted. Where possible butt welds will be radiographed to 
ensure no defects are present at the weld that could initiate damage during the experiments, as this is 
outside the scope of this project. 

The WPS (Welding Procedure Specifications), as well as typical material deliverables for the straight 
pipe sections, the Tee joints and the elbows are included in the corresponding section of deliverable D2. 

II.3.  Work  package  3  –  Experimental  testing  of  structural  components  of  critical 

importance under strong cyclic loading 

The work within WP3 consists of conducting a number of experiments in key components of industrial 
piping systems. The specimens are typical piping components from industrial facilities, compatible with 
the case studies examined in the first two work packages. Pipe diameter of the specimens is equal to 6 
and 8-inch, with standard thickness (SCH40) for all specimens. A brief summary of the tests is shown 
in  Table  II.3-1,  whereas  a  detailed  presentation  of  the  tests  is  offered  in  the  next  paragraphs.  For 
uniformity  the  material  of  all  specimens  is  P355N  according  to  EN  10216-3  standard  [21]  which  is 
equivalent to API 5L Grade X52. This choice has been based on the fact that this  steel grade is being 
currently considered for the design and construction of new industrial facilities, instead of lower steel 
grades. The P355 material is quite ductile and therefore, the test results can be also applicable to piping 
components  made  of  lower  steel  grade  (e.g.  P235,  P275).  Strong  cyclic  loading  has  been  applied  in 
order  to  examine  the  structural  behavior  of  the  components  under  repeated  excursions  well  into  the 
plastic range, simulating the loading conditions of a severe seismic event. 

25 
 
 
 
 
Table II.3-1: Summary of experimental tests 

Six  (6)  experiments  on  tank  openings 
(nozzles) – Task 3.2 

Fifteen  (15)  experiments  on  bolted 
(flanged) piping connections – Task 3.3 

Thirteen 
elbows – Task 3.4 

(13)  experiments  on  pipe 

Pseudo-dynamic 
a 
typical  piping  system  (4  levels)  –  Task 
3.4 

experiments  of 

Eight  (8)  tests  on  Tee-branch  piping 
connections – Task 3.5 

Testing of base steel material and welds 
- Task3.1 

Tests  on  6-inch  tank  openings  (nozzles)  of  3  different  shell 
nozzle  reinforcement  type,  under  2  different  load  directions 
(longitudinal and transverse) 
Tests  on  8-inch  SCH40  long-radius  induction  bends  (
1.5), connected to two straight pipe segments of 
Eight (8) tests on non-pressurized elbows and five (5) tests on 
pressurized elbows. 

 length. 

Pseudo-dynamic test of piping system with 8 and 6-inch pipes, 
3  elbows,  1  flange  and  1  Tee-  junction.  Four  (4)  levels  of 
earthquake excitation with internal pressure 3.2 MPa. 
Seven (7) bending tests (2 monotonic and 5 cyclic) have been 
conducted on pressurized standard joints (PN63, PN40).  
Eight bending and axial tests (3monotonic and 5 cyclic tests) 
have been also conducted on pressurized non-standard flanged 
joints.  
Eight (8) tests on  6-to-8-inch Tee-junctions specimens of two 
types  of  branch  to  main  steel  pipe  connections;  connection 
with  a  fitting  and  Tee-junction  with  a  welded  reinforcement 
plate. Four (4) tests on Tee fittings and four (4) tests on weld 
reinforced Tees. 
Base material characterization. 
Monotonic tensile testing for all pipe/plates. 
Cyclic loading (strain and stress controlled) on coupons, from 
elbows and plates.  
Charpy V-notch testing on elbow material. 
Low cycle fatigue tests of butt-welded coupon specimens. 
Weld characterization. 

II.3.1. 

Task 3.1: Basic material testing and characterization 

The  objective  of  this  task  was  to  investigate  the  behavior  of  the  base  material  of  piping  components 
selected for full scale testing and the properties of butt welded joints employed in construction of the 
full-scale  specimens.  The  results  obtained  by  the  characterization  campaign  have  been  introduced  as 
input in FE models (described in WP4) in order to simulate as realistically as possible both monotonic 
and cyclic loading of the piping components. Whenever suggested by the stress-strain field expected in 
full scale testing, specimens have been extracted in longitudinal and transverse direction. The following 
Table II.3.1-1 reports a summary of the small scale testing activity performed by EBETAM and CSM 
within  work  package  3,  task  3.1.  It  can  be  concluded  that  the  performed  material  characterization 
campaign is fully adequate with the aims of the project and the initial planned activity. Main results of 
this activity can be found in the Appendix III.3.1, while the whole set of results are extensively reported 
in the deliverable D3.1. 

RD5D26 
 
 
 
 
 
Table II.3.1-1: Small scale testing performed to support full scale activity and FEA 

Component 

Material 

test type 

Number of 
tests 

corresponding full-
scale testing 

8" x 8.18mm Elbow 

X52 
(equiv. to 
P355) 

8" x 8.18mm Pipe 

tensile-monotonic 

cyclic/imposed strain 

cyclic/imposed stress 

charpy V-notch 

tensile-monotonic 

8" x 8.18mm Pipe 

P355 

tensile-monotonic 

20mm Plate 

P355 

8"x8.18mm Pipe 

6"x7.1 mm Pipe 

8"x6" Fitting 

P355 

8"x8.18mm Pipe, 
butt weld joint 

P355 

tensile-monotonic 

cyclic/imposed strain 

tensile-monotonic 

Ring test monotonic 

tensile-monotonic 

tensile-monotonic 

magnetic particle inspection 

radiographic inspection 

macroscopic examination 

hardness tests 

bend test 

transverse tensile test 

charpy v-notch 

cyclic/imposed stress 

4 

6 

6 

8 

4 

4 

4 

3 

2 

2 

2 

2 

1 

1 

1 

1 

4 

2 

6 

9 

Elbows 

Flanges  

Nozzles  

Tee joints 

Validation of WPS 
employed in full 
scale testing 

II.3.2. 

Task 3.2: Tests on tank openings/nozzles under strong cyclic loading 

The test program comprised in total 6 cyclic tests with 3 different shell nozzle reinforcement types and 
2 different load directions (longitudinal and transverse): 

 

 

 

P1: Nozzle with internal reinforcing plate 

P2: Nozzle with external reinforcing plate 

P3: Thickened nozzle reinforcement 

The 6 specimens have been manufactured by a company which is specialized in manufacturing of flat-
bottomed steel tanks according to the state-of-the-art. Different examples for flat-bottomed steel tanks 
have  been  compiled  from  literature  as  well  as  from  direct  dialogue  with  manufacturers  to  mirror  the 
state-of-the-art. Following these examples one reference example of a complete tank has been agreed on 
and the necessary size (cut-out from the complete tank) of the test specimens has been determined using 
numerical  calculations  by comparing  the stresses  and  deformations  at  the  nozzle  and the  surrounding 
area of the complete tank with those at the nozzle of the test specimen. Further on plate thickness and 
joint details have been assigned according to the codes provisions of EN 14015 [26]. 

For  the  test  setup  the  boundary  conditions  representing  the  cut-out  from  the  complete  tank  were 
designed and realized in order to represent in a best possible way the real load-deformation behaviour 
of  the  complete  tank.  The  most  critical  parameter  so  far  was  assumingly  the  possibility  of  the  tank 
bottom  plate  to  lift  up.  Free  rotation  of  the  flanges  was  allowed  for  by  a  load  application  using 
pendulum rods. The final complete test setup with test specimen 3-1 can be seen in Figure II.3.2-1. 

27 
 
Figure II.3.2-1: Test setup in the laboratory with test specimen P3-1. 

According to the ECCS-recommendation [20] monotonic tests are required to define the test procedure 
for cyclic testing. Instead of these monotonic tests, here numerical simulations were used. The results of 
these  simulations  are  given  in  III.3.2.  Since  the  displacements  ey  are  similar  for  all  3  reinforcement 
types, the average values ey = 10 mm for all tests with longitudinal load direction and ey = 20 mm for all 
tests  with  transverse  load  direction  were  taken  for  the  laboratory  cyclic  tests.  This  procedure  also 
allowed for a good comparison of the results. The tests were run displacement-controlled, whereby the 
loading was applied as given in Figure II.3.2-2 in accordance to ECCS [20]. 

Figure II.3.2-2: Schematic representation of the loading history according to ECCS [20] 

The test results in terms of cyclic load-displacement curves and pictures of the specimens at the end of 
the loading history are shown in  III.3.2. Table II.3.2-1 summarized the number of cycles until failure 
detected during the test execution. 

Table II.3.2-1: Number of cycles until failure Nf for longitudinal and transverse load direction 

Specimen 

P1-1 Internal 

load 
direction 

P2-1 External 

longitudinal 

P3-1 Nozzle 

P1-2 Internal 

P2-2 External 

transverse 

P3-2 Nozzle 

Nf 

24 
24 
16 
9 
9 
11 

28 
 
 
 
 
 
 
 
 
II.3.3. 

Task 3.3: Tests on bolted piping connections under strong cyclic loading 

II.3.3.1.

Tests on standard bolted piping connections under strong cyclic loading 

CSM performed a total of seven full scale tests on standard P355N steel grade circular bolted flanged 
connections  according  to  Euro  code  EN  1092-1  [23],  Figure  II.3.3-1.Two  types  of  flanges  have  been 
tested: PN63 DN200 and PN40 DN 200 weld neck flanges, type 11. The aim of the testing activity was 
to investigate the behavior of P355N steel piping flanged joint both in monotonic than in cyclic bending 
loading,  with  internal  pressure.  Testing  have  been  performed  adopting  the  4-point  bending 
configuration, Figure II.3.3-2, in order to guarantee constant bending moment over the test section, thus 
allowing to study the critical parts of the connection under the same load conditions. Internal pressure 
has  been  obtained  by  means  of  water,  and  the  axial  force  due  to  pressure  acting  on  pipe  end  caps  is 
developed. A picture of the test device that has been extensively modified by CSM on for the present 
test campaign is reported in Figure II.3.3-3.Two bending displacement amplitudes have been adopted: 
two times and four times the displacement causing yielding of the pipe section. Calculated according 
the procedure reported in the ECCS no 45 loading protocol [20], and two levels of internal pressure: 50 
and 32 bars equal to 80% of nominal connection operative pressure.  

Figure II.3.3-1: Weld neck flange geometry 

Figure II.3.3-2: Test configuration, four point bending 

Figure II.3.3-3: Testing device 

The test program on standard bolted flanged joints and main results / observations are reported in Table 
II.3.1-1.  Load  vs.  stroke  actuator  diagram,  and  pictures  of the  specimens  after  testing  are  reported  in 
III.3.3.1. 

Geometry 1: PN63/DN200, material P3553 tests:1 monotonic axial 2 bending + constant axial forceGeometry 2: PN40/DN200, material P3553 tests:1 monotonic axial 2 bending + constant axial force29 
 
 
 
 
 
 
 
 
 
 
Table II.3.3-1: CSM test program on standard bolted flanged joints and main results / observations 

Test 

Test type 

Leakage 

Load 

Cycles 
to 
failure 

PN63-1 

Monotonic 

157 kNm 

- 

147 kNm 

26th 

PN63-2 

PN63-3 

Constant 
Amplitude 

=+/-54mm 

Constant 
Amplitude 
=+/-27mm 

Observations 

Pipe 

Flange 

Gasket 

Bolts 

Welding 

Plastic 
deformation 

Wrinkling/ 

cracking 

No  

deformation 

Little plastic 
deformation 

No deformation 

No 
deformation 

Little plastic 
deformation 

No deformation 

144 kNm 

323th 

Cracking 

No 
deformation 

Crush due to 
compression 

No deformation 

PN40-1 

Monotonic 

109 kNm 

- 

Buckling 

PN40-2 

PN40-3 

PN40-4 
(extra) 

Constant 
Amplitude 
=+/-54mm 

Constant 
Amplitude 
=+/-27mm 

Constant 
Amplitude 
=+/-54mm 

103 kNm 

7th 

104 kNm 

232th 

110 kNm 

6th 

Buckling/ 

cracking 

Wrinkling/ 

cracking 

Buckling/ 

cracking 

No 
deformation 

Little plastic 
deformation 

No deformation 

No 
deformation 

Little plastic 
deformation 

No deformation 

No 
deformation 

Little plastic 
deformation 

No deformation 

No 
deformation 

No 
deformation 

Little plastic 
deformation 

No deformation 

No 
deformation 

No 
deformation 

No 
deformation 

No 
deformation 

No 
deformation 

No 
deformation 

Large  plastic  deformation of  the  pipes is  observed,  wrinkling  or local  buckling  is  developed in  cases 
were bending amplitude =+/-54mm is applied. The welds were not affected by cracking in any of the 
analyzed  cases.  Flanged  joints  shown  high  resistance  to  bending  loading  ensuring  leak  tightness  far 
beyond the pipe elastic limit, and no damage to bolts or flange at final unloading. It is thus concluded 
that Standard welding neck joints are suitable to be employed in operating conditions where a seismic 
event causes pipe bending present. On the other hand the life of the join is influenced by the  pipe that 
has to be carefully taken into account. 

II.3.3.2.

Tests on non-standard bolted piping connections under strong cyclic loading 

An  experimental  test  campaign  on  some  non-standard  Bolted  Flange  Joints  was  conducted  by  the 
University of Trento (UTrento) in order to assess their capacity and performance under seismic loading. 
A detail report on the experimental activities can be found in Deliverable D3. Although Bolted Flange 
Joints  (BFJs)  -one  of  the  most  critical  components  of  a  piping  system-  are  vulnerable  to  earthquake 
loading, they do not have any design Standard that takes into account seismic loading. Current design 
Standards,  e.g.,  EN1092-1  [23],  EN1591-1  [28],  ASME  B16.5  [6],  mainly  ensure  joint  integrity  and 
leak tightness under operating conditions but their high thicknesses make them unsuitable for seismic 
applications. In this respect, two non-standard BFJs -Design 01 and Design 02- comparatively thinner 
than  the  Standard  ones  were  designed  by  UTrento  to  achieve  a  better  performance  under  seismic 
loading. Initial dimensions of the flanges were taken from Eurocode EN 1092-1 [23] to be used with an 
8” connecting pipe under regular operating conditions of a petrochemical industry. The thickness of this 
standard  flange  was  then  reduced  based  on  mode  #1  and  mode  #2  failures  according  to  structural 
Eurocode EN 1993-1-8 [29]. Dimensions of other components of the BFJs were selected from relevant 
European  standards.  Materials  compatible  with  design  operating  conditions  were  chosen  for  different 
components  of  the  flange  joints.  A  calculation  on  the  designed  BFJs  according  to  the  new  European 
standard  EN  1591-1  [28]  was  performed  to  check  their  mechanical  integrity  and  leak-tightness 
performance.  Dimensions  of  the  non-standard  BFJs  are  presented  in  Figure  III.3.3-17  and  Table 
III.3.3-1 (all figures and tables referred in this Section are presented in III.3.3.2). 

A number of bending and axial experimental tests as listed in  Table III.3.3-2 were carried out on the 
designed non-standard BFJs under monotonic and cyclic loading in order to investigate their capacity 
and performance. A moderate internal pressure of 1.5 MPa was used in all tests. The monotonic tests 
were mainly performed to find required parameters to build ECCS45 [20] loading protocols to be used 
for the cyclic tests. Test specimens were constructed by placing the flange joints in the middle of two 

30 
 
 
 
 
connecting  pipes  as  depicted  in  Figure  III.3.3-18  of  III.3.3.2.  Loading  were  applied  to  the  test 
specimens  by  means  of  MOOG  actuators  with  a  capacity  of  1000  kN.  During  bending  tests,  loading 
were applied till the failure of the test specimens, whereas, the maximum level of loading during the 
axial  tests  was  limited  by  the  capacity  of  the  actuators.  The  experimental  set-ups  are  illustrated  in 
Figure  III.3.3-19  of  III.3.3.2.  Experimental  results  were  compared  with  the  allowable,  yield  and 
ultimate loads suggested by codes. A comparison was also made between the capacity of the designed 
joints and the seismic demands on such joints coming from a case study performed on a petrochemical 
piping system. 

Both  of  the  non-standard  BFJs  exhibited  favourable  performance  under  bending  and  axial  loading  as 
can be noted from Figure III.3.3-20 and Figure III.3.3-21 of III.3.3.2. A good capacity of the joints in 
terms  of  strength,  ductility  and  energy  dissipation  were  found  in  all  experimental  tests.  Almost  no 
degradation of the joints in cyclic loading was observed and no failure occurred in any of the joints. In 
all tests, leakage was observed with a high level of loading. During bending tests, failure took place in 
the  pipe  wall  where  local  buckling  was  observed  as  shown  in  Figure  III.3.3-22.  During  axial  tests, 
loading  were  applied  by  means  of  two  MOOG  actuators  each  with  a  capacity  of  1000  kN.  Hence,  a 
maximum axial load of 2000 kN could be applied during axial tests. With this level of maximum load, 
test  specimens  remained  within  its  elastic  limit  while  only  small  levels  of  flange  displacement  were 
found. To  assess  low  cycle  fatigue behaviour  of the BFJs  in  axial  loading,  once  the  limit load of the 
actuators  was  reached,  we  continued  several  cycles  of  loading  with  the  maximum  displacement 
attained. Nevertheless, no failure or yielding occurred in the flanges or in any part of the joints.  

The comparison between experimental results and Coded loads displayed a favourable performance of 
the  non-standard  BFJs.  Allowable  seismic  design  loads  calculated  according  to  Codes  were 
considerably lower than experimental leakage and yielding loads as can be noted from Table III.3.3-3 – 
Table  III.3.3-4.  The  levels  of  earthquake  demands  on  a  piping  system  found  from  a  case  study,  see 
Table III.3.3-5, were significantly lower than the capacity of the designed BFJs. Thus, the non-standard 
BFJs were considered suitable for regular seismic applications. It was recommended that these types of 
joints  could  be  used  in  piping  systems  operating  both  under  normal  conditions  and  under  seismic 
events. 

II.3.4. 

Task 3.4: Tests on piping elbows under strong cyclic loading 

II.3.4.1.

Tests on non-pressurized piping elbows under strong cyclic loading 

219.1  mm  and 

Eight tests have been conducted on 8-inch SCH40 long-radius induction bends (nominal diameter and 
1.5),  connected  on  two 
thickness  equal  to 
straight  pipe  segments  with  length  equal  to 
and  the  same  nominal  diameter  and  thickness.  The 
material of the specimens is P355N according to EN 10216 standard [21], equivalent of API 5L X52. 
The  uniaxial  and  cyclic  steel  material  properties  have  been  examined  in  detail  through  an  extensive 
experimental program [13].  

8.18  mm  respectively,  and 

The  experimental  campaign  took  place  at  Delft  University  of  Technology.  The  experimental  set-up 
consists  of  a  steel  frame  with  a  hydraulic  cylinder  positioned  horizontally  over  the  lower  part  of  the 
frame.  The  cylinder’s  one  end  is  pin-connected  on  the  frame  while  the  other  end  is  used  for  the 
application of end-displacements on the elbow. The elbow specimen is pin-supported on the same frame 
through a pair of steel plates attached to the end of the specimen. The hydraulic cylinder is connected to 
the other end of the specimen using a similar support, as shown in  Figure II.3.4-1. Strain gauges  and 
displacement  transducers  are  used  to  monitor  the  evolution  of  local  strains  at  critical  location  of  the 
specimens  and  the  displacements  of  the  elbow  ends  respectively  (Figure  II.3.4-2).  The  change  of  the 
elbow  vertical  diameter  is  monitored  at  its  middle  cross-section  and  at  a  cross-section  near  the  weld 
region using two special-purpose frame devices. Thickness measurements on the specimens have been 
obtained prior the execution of every test and differences with respect to the nominal value up to 1.9 
mm (23.3 %) were recorded. The specimens are filled with water before testing with low water pressure 
equal to 1.2 bar (0.5% of the yield pressure), which is kept constant throughout the test and used as an 
indicator of through-thickness crack opening. 

DtRD5D31 
 
 
 
Figure II.3.4-1: Experimental set-up: Front view  

Figure II.3.4-2: Experimental set-up: Instrumentation 

Test 
No. 

Displacement 
(mm) 
range 

Number of cycles 
until failure 

Local strain 
range* (%) 

1 

2 

3 

4 

5 

6 

7 

8 

±25 

±70 

±100 

±150 

±200 

±250 

±300 

ECCS protocol 

13160 

444 

171 

61 

28 

17 

10 

16 

(Exp.) 

0.33 

1.23 

- 

2.61 

- 

3.84 

4.02 

- 

*: calculated in the elbow hoop direction at the critical 
location 

Table II.3.4-1: Displacement range and number of cycles 

Figure 
configuration 

II.3.4-3:  Cracked 

elbow 

specimen 

The specimens have been subjected to intense cyclic in-plane bending through the application of cyclic 
end-displacements  on  the  moving  support,  keeping  the  other  support  pin-connected  to  the  supporting 
steel frame. The seven first tests followed a constant amplitude end-displacement, whereas a variable-
amplitude loading sequence has been applied to the eighth specimen. The number of cycles until failure 
correspond  to  the  development  of 
of  each  specimen  is  tabulated  in  Table  II.3.4-1.  The  values  of 

through-thickness crack  and  leakage  of  containment, as  shown  in  Figure  II.3.4-3.  It  is  noted that this 
failure  pattern  is  similar  for  all  specimens.  In  Figure  II.3.4-4  a  summary  of  the  stabilized  load-
displacement curves for the five tests conducted is presented. 

The eighth test has been conducted following an increasing end-displacement pattern, shown in Figure 
III.3.4-3  (Appendix  III.3.4.1),  according  to  the  loading  protocol  for  cyclic  loading  proposed  by  the 
ECCS Recommendations for cyclic testing of structural components  [20]. Note that in the case of an 
elbow,  because  of  its  significantly  different  behavior  under  closing  and  opening  bending,  the 
displacement pattern is not symmetric, shown in  Figure II.3.4-5. Using this end-displacement pattern, 
the specimen has failed in the form of fatigue cracking after 16 completed loading cycles. 

lfNfN32 
 
 
 
 
 
 
 
Figure  II.3.4-4:  First  load-displacement  cycles  for 
the 5 tests 

Figure II.3.4-5: Test 8 – ECCS load - displacement 
cycles 

II.3.4.2.

Tests on pressurized piping elbows under strong cyclic loading 

Five full scale tests on API 5L X52 grade, 8” piping sections containing an elbow connected by welding 
to two straight pipes and subjected to cyclic bending moment and internal pressure. The geometry of the 
specimen  is  reported  in  Figure  II.3.4-6  and  in  Table  II.3.4-2,  while  Table  II.3.4-3  reports  the  test 
program performed. One extra test with respect to the original test plane has been performed adopting a 
spare specimen available. The experimental activity has been performed adopting the CSM Full scale 
test machine for offshore tubular components, located in the CSM laboratory in Sardinia. A picture of a 
specimen installed in the experimental device is reported in Figure II.3.4-7. 
Table II.3.4-4 summarized the main outcome of the test campaign. All specimens have been tested up to 
final failure, occurred by through wall cracking at the elbow mid-section, about 20mm below the flank 
line, Figure II.3.4-8. Relevant ovalization occurred in all tests on the elbow cross section, which caused 
large  plastic  strain accumulation.  Results  of Tests  1  to  4  show that the  displacement  amplitude has a 
marked effect on the number of cycles to failure, while the internal pressure level, raise from 32 to 70 
bar has a weaker effect. A further increase of internal pressure to 120 bar with a displacement amplitude 
of +/-200mm (test 5) resulted in a reduction of cycles to failure form 27 to 22. For all levels of internal 
pressure and displacement amplitude, although large bending displacement has been applied, no local 
buckling has been observed either on pipes and elbows before to cracking occurrence. No cracking is 
observed on welds to the straight pipe segments or in other locations. 

D  =  219.1  mm  8.625 

inch 

WT  =  8.18  mm  0.322 

inch 

R/D  =  1.5 

– 

– 

inch 

L  =  1100  mm 

Grade  =  API 
5L 

X52 

Figure II.3.4-6: Specimen geometry and test configuration  Table II.3.4-2: Specimen properties  

33 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
Table II.3.4-3: Full scale test program 

Test ID 

End displacement (mm)  Δδ (mm) 

1st cycle 

1 

2 

3 

4 

5 (extra test) 

±200 

±200 

±300 

±300 

±200 

Internal 
Pressure 

3.2 MPa 

7.0 MPa 

400 

400 

600 

Closing mode 

3.2 MPa 

600 

400 

7.0 MPa 

10-12 MPa 

Figure  II.3.4-7:  Specimen  inside  the  experimental 
device 

Figure  II.3.4-8:  Specimen  failure  for  cracking  at  flank 
position 

Test 
ID 

Applied 
displacement 

Internal 
Pressure 

Cycles at 
failure 

1 

2 

3 

4 

5 

±200 mm 

3.2 MPa 

±200 mm 

7.0 MPa 

±300 mm 

3.2 MPa 

±300 mm 

7.0 MPa 

±200 mm 

12 MPa 

26th 

27th 

10th 

10th 

22th 

Crack 
position 

below flank 

17 mm  

15 mm  

18 mm  

24 mm  

20 mm  

Table  II.3.4-4:  Summary  of  Elbows  test  results  performed  by 
CSM 

Figure  II.3.4-9:  Typical  Force  vs.  displacement 
result, Test 2 

II.3.4.3.

Pseudo-dynamic loading on piping system 

The University of Trento (UTrento) undertook a pseudo-dynamic test campaign on a typical full-scale 
piping  system  containing  several  critical  components  in  order  to  assess  its  seismic  performance. 
Pseudo-Dynamic testing with Dynamic substructuring (PDDS) is a novel hybrid experimental technique 
in  which  a  heterogeneous  model  of  the  emulated  system  is  created  by  combining  a  Physical 
Substructure (PS) with a Numerical Substructure (NS) that describes the remainder of the system. The 
overall response of a structure is evaluated by combining the experimental response of the PS  -which is 
generally the most critical part of the structure- with the numerical response of the NS. By the use of 
actuators,  computers,  finite  element  software  and  controllers,  this  tests  can  potentially  reproduce  a 
shaking table test. A number of PDDS as reported in Table III.3.4-2 (all figures and tables referred in 

34 
 
 
 
 
 
 
 
 
 
 
this section are presented in  III.3.4.3) were carried out on the piping system under different levels of 
earthquake loading corresponding to serviceability and ultimate limit states suggested by Standards. 

The  piping  system  experimentally  tested  was  a  typical  industrial  piping  system  placed  on  a  support 
structure as shown in Figure III.3.4-24(a). It contained 8” and 6” straight pipes, several elbows, a Tee 
joint and an EN 1092-1 [23] Standard PN 40 weld-neck bolted flange joint. Material of all straight pipes 
and elbow elements was API 5L Gr. X52. The pipes were filled with water with an internal pressure of 
3.2 MPa. Preliminary numerical analyses were carried out through a 3D Finite Element (FE) model of 
the  piping  system  developed  in  SAP2000  [50]  software.  The  inputs  earthquakes  corresponding  to 
serviceability and ultimate limit states suggested by Standards, see Table III.3.4-3, were found on the 
elevated floor of the support structure through time history analyses by imposing an earthquake to its 
base.  An  elastic  seismic  analysis  was  carried  out  on  the  piping  system  with  and  without  taking  into 
account  a  Rayleigh  damping  constructed  considering  a  0.5%  damping  found  through  Identification 
Tests (IDTs) on the PS. The Rayleigh’s damping considerably reduced the stress level as can be noted 
from Table III.3.4-4. 

Implementation of the PDDS algorithm on the piping system was challenging mainly because it was a 
structure  with  distributed  mass  –for  which  the  PDDS  has  been  considered  inadequate–  and  it  was 
subjected to distributed earthquake forces. Two coupling points were chosen to divide the piping system 
into a PS and NS based on the locations of bending moments close to zero in the x-y plane, see Figure 
III.3.4-24(b), found from a time history analysis. Since the PDDS were carried out at an elongated time-
scale  compared  to the  real-time  of  an  earthquake,  the  inertia and  damping  forces  had  to be  modelled 
numerically. Additionally, earthquake forces were to be condensed to the two coupling nodes since it 
was not feasible to apply an actuator in each node of the PS to produce the earthquake forces. In order 
to condense the PS to the two coupling nodes, the so called Craig-Bampton (CB) [18] mode synthesis 
or model reduction technique –which was found to be the most effective among others- was adopted. 
The  PDDS  test  algorithm  was  developed  in  Matlab  and  Simulink  using  the  LSRT2  integration 
algorithm developed by Bursi et al. [9]. Experiments were carried out using an xPC target, two MOOG 
actuators and an MTS FT60 controller [44]. The experimental set-up was realized in the Materials and 
Structural  testing  laboratory  of  UTrento.  The  schematic  and  actual  test  set-up  is  presented  in  Figure 
III.3.4-25 and Figure III.3.4-26, respectively, whereas, some components of the PS are shown in Figure 
III.3.4-27. 

Several PDDS as listed in Table III.3.4-2 were carried out on the piping system under different levels of 
earthquake  PGAs  corresponding  to  serviceability  and  ultimate  limit  states  prescribed  by  an  Italian 
Performance-Based  Earthquake  Engineering  Standard  as  listed  in  Table  III.3.4-3.  Some  IDTs  were 
carried  out  to  estimate  the  damping  to  be  used  during  experimental  tests.  A  favourable  seismic 
performance of the piping system was observed during the experiments. It was found that, even under 
the Collapse Limit State level earthquake, both the piping system and its critical components did not fail 
or yield. Small levels of rotations and strains were found in different components; see Figure III.3.4-28 
and Figure III.3.4-29. In all tests, a good agreement between the experimental and numerical results was 
observed  as  can  be  noted  from  Figure  III.3.4-30  in  this  respect.  Note  that  such  behaviour  of  piping 
system  is  desirable.  In  fact,  many  researchers  reported  similar  performance  of  piping  systems  under 
earthquake loading based on shaking table tests. Therefore, we concluded that present seismic design 
rules for piping systems and components are conservative and amendments are necessary to overcome 
some degree of conservatism. 

II.3.5. 

Task 3.5: Tests on T-branch pipe connections under strong cyclic loading 

This  experimental  investigation  consists  of  eight  (8)  tests  on  T-joint  pipe  connections.  Two  types  of 
branch  to  main  pipe  connections  are  considered:  connection  with  a  fitting,  denoted  as  TJF  (Figure 
II.3.5-1a) and T-joint with a welded reinforcement plate, denoted as TJWR (Figure II.3.5-1b). The pipes 
are of P355NH/NL1 steel (Grade X52) with a yield stress of 412 MPa and the steel fitting has a nominal 
yield stress of 355 MPa. The specimens have been subjected to monotonic and low cycle fatigue tests 
until  a  through-crack  failure  occurred.  For  the  fully-reversing  cyclic  tests  three  (3)  imposed 
displacement  histories  were  chosen,  two  with  constant  displacement  amplitude  and  one  according  to 
ECCS No45 [20] testing procedure. The test results are presented in Table II.3.5-1. 

35 
 
The branch pipe has a nominal cross-section 168.3x7.1 mm and the main pipe 219.1x8.2 mm. The same 
setup arrangement was used for both the static monotonic and cyclic tests, as shown in Figure II.3.5-2. 
The load was applied at the free-end of the branch through a special ball-joint hinge, as shown in Figure 
II.3.5-2b. Both ends of the main pipe of the specimens were bolted on the testing frame. The specimens 
were filled with air (0.5 bar pressure) in order to determine the moment that the through-thickness crack 
occurred. For all the T-joint tests, strain gages were placed at critical locations to determine the strain 
distribution in the region of the joint studied. In addition, wire position transducers were used for the 
load-point displacement (LPD) measurements. 

Table II.3.5-1: Test results. 

type of 
specimen 

TJF 

TJWR 

1M 
2C
WR 
3C
WR 
4C
WR 
1M
WR 
2C
WR 
3C
WR 
4C
WR 

type of static 
loading 
monotonic 
cyclic 
cyclic 
cyclic 
monotonic 
cyclic 
cyclic 
cyclic 

peak load, 
Pmax (kN) 
158.0(*) 
131.2 
117.9 
159.9 
81.9(*) 
78.3 
64.9 
77.9 

displacement,  
δ (mm) 
200.0 
60 
45 
ECCS No45 
200.0 
60 
45 
ECCS No45 

loading cycles 
to failure, Nf 
- 
42 
135 
15 
- 
6 
17 
11 

(*) load corresponding to an imposed LPD of 200 mm 

                                (a)            

      (b) 

Figure  II.3.5-1:  The  two  types  of  T-joint  branch  connection  studied:  (a)  fitting,  (b)  welded  reinforcement  ring-
plate. 

Figure II.3.5-2:  Test setup for the T-joint tests: (a) overall view, (b) side view. 

       (a)            

 (b) 

36 
 
 
 
 
 
 
 
 
 
                             
 
                              
 
(a) TJF (fitting) tests: one specimen was subjected to static monotonic loading and three (3) to low-
cycle  fatigue.  For  an  imposed  LPD  of  200  mm,  the  corresponding  carrying  capacity  of  the  specimen 
was  158.0  kN.  The  through-thickness  crack  occurred  in  the  region  of  the  fitting  of  the  T-joint 
connection (see  deformed specimen  in  Figure  II.3.5-3a). Two specimens  were subjected to low-cycle 
fatigue  under  displacement-control  at  an  imposed  LPD  value  of  δ1=60mm,  δ2=45mm  and  one 
specimen to a cyclic loading history according to the ECCS No45 testing procedure, at a frequency of 
0.010.02 Hz. The specimens failed after Nf = 42, 135 and 15 cycles, respectively. All specimens failed 
when a through-thickness crack occurred (at the same location as in the test under monotonic load). 

(b) TJWR (welded reinforcement) tests: one specimen was subjected to static monotonic loading and 
three  (3)  specimens  to  low-cycle  fatigue.  For  an  imposed  LPD  of  200  mm,  the  corresponding  load 
carrying  capacity  of  the  specimen  was  81.9  kN.  The  failure  occurred  in  the  branch  region  of  the 
connection (see deformed specimen in  Figure  II.3.5-3b). Two specimens were subjected to low-cycle 
fatigue  under  displacement-control  at  an  imposed  LPD  value  of  δ1=60mm,  δ2=45mm  and  one 
specimen  to  a  cyclic  loading  history  according  to  ECCS  No45  testing  procedure,  at  a  frequency  of 
0.010.02 Hz. The specimen tested at an imposed displacement of 60 mm, failed after Nf= 6 cycles at 
the same location as in the monotonic test. The other two specimens failed after Nf=17 and 11 cycles, 
respectively, in the welded reinforcement plate region of the connection. 

Figure II.3.5-3: Deformed T-joint specimens at failure: (a) TJF, (b) TJWR. 

   (a)                                                                  

 (b)                                       

II.4.  Work  package  4  –  Numerical  simulation  of  industrial  components  behaviour 

under strong cyclic loading 

The objectives of WP4 are: 

•  The definition and calibration of an appropriate cyclic-loading constitutive material model to use 
for  a  realistic  simulation  of  elasto-plastic  behaviour  of  tubular  components;  the  constitutive 
material  model  should  be  able  to  simulate  material  degradation  effects,  as  well  as  low-cycle 
fatigue.  

•  The development of a modeling finite element procedure that simulates accurately the behaviour of 
industrial components; the modeling procedure will be validated through the experimental data of 
the previous workpackage (WP3). 

•  The  performance  of  an  extensive  parametric  study,  using  the  numerical  modeling  methodology 
developed  above,  to  examine  the  behaviour  of  structural  components,  within  a  wide  range  of 
geometric and material parameters 

II.4.1. 

Task 4.1 Definition of inelastic material modelling under cyclic loading 

II.4.1.1.

Introduction 

The  behavior  of  metal  structures  that  undergo  plastic  deformations  can  be  effectively  described 
adopting a suitable plasticity model. Several models are available nowadays in the relevant literature. 
Some  of  these  are  already  implemented  in  commercial  finite  element  codes,  while  others  require  the 

37 
 
 
       
 
 
 
 
 
development  of  user-subroutines  for  their  implementation  and  use.  In  the  following  paragraphs  the 
proposed plasticity models for the simulation of cyclic loading of metals are presented. These advanced 
models have been selected over other more simplified models (e.g. linear kinematic hardening model) 
because of their ability to describe more accurately the Bauschinger effect, which is evident upon load 
reversal of the material stress-strain curve in the plastic range and may affect the overall behavior of the 
elbow specimens. 

In the present study, the following models will be used (a) the Armstrong-Frederick model and (b) the 
Tseng-Lee  model.  In  most  of  the  cases,  the  Armstrong-Frederick  model  will  be  employed. 
Nevertheless, in a few cases, the Tseng-Lee model will be used, mainly for comparison purposes. 

II.4.1.2.

The Armstrong-Frederick model and calibration 

The  Armstrong-Frederick  (A-F)  model  has  been  selected  among  other  models  to  be  used  in  the 
numerical  simulations  within  the  INDUSE  project.  It  is  an  advanced  plasticity  model,  significantly 
more powerful than linear kinematic models, available in many commercial finite element codes (e.g. 
ABAQUS, MARK etc.). Within J2-plasticity framework, the Armstrong  - Frederick model in its most 
general version allows the size of the yield surface to change together with translation. The equation for 
the yield surface is written as: 

(1) 

where 

is  the  deviatoric  stress tensor, 

  is the  so-called  “backstress”  tensor  and 

is  a function 

that defines the size of the yield surface. In general 

can be a function of the equivalent plastic 

strain 

 or it can be assumed to be constant, as in the present case.  The model assumes a nonlinear 

kinematic hardening rule for the evolution of the back stress written in the form: 

(2) 

where 

and 

are model parameters that need to be calibrated. 

The calibration procedure is based on material test results obtained by CSM. The coupon specimens for 
the  material  testing  were  extracted  from  the  longitudinal  and  hoop  direction  elbow  of  specimens 
planned to be tested under cyclic loading conditions.  The behavior of the elbows is governed mainly by 
the material properties in the hoop direction, therefore the corresponding test results have been used for 
and .  A  summary  of  these  tests  is  presented  in 
the  calibration  of  the  material  model  parameters 
Figure II.4.1-1. 

Lemaitre and Chaboche have proposed a systematic calibration method based on cyclic stain controlled 
material  tests,  as  those  proposed  in  Figure  II.4.1-2.  According  to  this  method,  by  using  the  three 
 is defined. Subsequently, for each 
available test stress-strain curves, a mean size of the yield surface 
point is defined and the points are plotted in a graph. In the same graph a 
test curve, one

curve is fitted following the equation:  

(3) 

Parameters 
II.4.1-2). The resulting 

and 

 are selected to provide the best fit of the curve to the experimental points (Figure 

and 

 by this method are reported as Set 1 in Table II.4.1-1. 

2()1()()023sasaqkFsa()qk()qkqaεapqCCCkpΔσΔε22-k,pΔσCΔε-k=tanhγ2γ2CC38 
 
 
 
 
 
 
 
 
 
Figure  II.4.1-1:  Material  test  results  –  elbow  hoop 
direction 

Figure II.4.1-2: Lemaitre-Chaboche calibration method 

Figure II.4.1-3: Calibration results – Test E1B T-4 

Figure II.4.1-4: Calibration results – Test E1B T-9 

Figure II.4.1-5: Calibration results – Test E1B T-5 

Figure  II.4.1-6:  Test  on  elbows  –  Simulation  results 
using Set 1 and Set 3 

Alternatively,  the  A-F  model  can  be  calibrated  to  fit  each  material  cyclic  stress-strain  curve.  The 
resulting parameter sets by this approach are denoted as Set 2, Set 3 and Set 4 for the tests E1B-T-4, 
E1B-T-9  and  E1B-T-5  respectively  (Figure  II.4.1-3  to  Figure  II.4.1-5).  Based  on  the  expected  strain 
range at the critical locations of the examined elements, the appropriate A-F model parameters (those 
taken from tests at similar strain levels) should be selected. Nevertheless, as it has been proven by the 
simulation results on elbows presented in Figure II.4.1-6, the adoption of different parameter sets does 
not affect the simulation accuracy significantly. Therefore, for simplicity the parameter Set 3 could be 
used for all the simulations. 

39 
 
 
 
 
 
 
 
 
 
Table II.4.1-1: Material model parameter values 

E (Mpa) 
210000  

Set 1  
Set 2  
Set 3  
Set 4  

ν  
0.3  
σy (MPa)  
317.83  
300  
300  
300  

C  
18000  
41000  
46000  
46000  

γ  
75  
200  
215  
185  

II.4.1.3.

Tseng-Lee model and calibration 

In the classical plasticity models as those presented in the previous paragraphs, only one surface is used, 
namely the yield surface (YS), for the determination of the limits of the elastic regime. Based on this 
formulation, when plastic loading takes place, the hardening modulus is defined through the consistency 
condition.  As  an alternative  to  this  approach, the “Bounding  Surface” concept  can  be  adopted.  In  the 
plasticity models following this concept, the hardening modulus is defined by a given expression, not 
through the consistency condition. 

The Tseng-Lee model [53] is an effective model to describe loading histories including cyclic plasticity 
adopting  the  so-called  “Bounding  Surface”  concept.  In  this  model,  in  addition  to  the  yield  surface,  a 
“memory surface” is introduced, which plays the role of a bound and obeys isotropic hardening (Figure 
II.4.1-7).  The  yield  (inner)  surface  (Y.S.)  is  free  to  translate  and  change  shape  within  the  memory 
surface  (M.S.)  through  a  mixed  (combined)  hardening  rule.  The  memory  surface  is  centered  at  the 
origin  and  hardens  only  isotropically  every  time  its  stress  level  is  exceeded.  Thus,  it  represents  the 
highest level of stress developed in the loading history. 

During  initial  plastic  loading,  the  flow  rule  is  based  on  the  yield  surface  and  the  hardening  modulus 
depends on the relative distance of the current stress on the yield surface and an appropriately chosen 
stress point on the memory surface. When the yield surface reaches the memory surface during further 
loading,  it  stays  attached  to  the  memory  surface  at  the  specific  stress  point.  The  two  surfaces  lose 
contact  when  the  first  reverse  loading  occurs  associated  with  plastic  deformation.  The  expressions 
describing the yield surface 

 and the memory surface 

 are respectively: 

(4) 

(5) 

 is the deviatoric part of the stress tensor 

where 
, 
is  the  deviatoric  stress  on  the  memory  surface  and 

is the deviatoric part of backstress tensor 

, 

,

  are  functions  of  equivalent  plastic 

strain

, representing the size of the yield and the memory surface respectively. 

In this model, similar to all uncoupled models, the hardening modulus 
is defined directly through an 
appropriate function. Motivated by experimental observations this definition of plastic modulus should 
account for smooth transition from the elastic to the inelastic stage. The model accounts for the gradual 
decrease of the value of 
as hardening proceeds, which is of particular importance for the successful 
modeling  of  strong  low-cycle  loading  conditions  associated  with  reverse  plastic  loading  and  the 
Bauschinger effect. 

The basic concept of this model is that the value of the hardening modulus 

 in stress space between the yield surface and the bounding surface. More specifically 

as follows: 

(6) 

depends on the "distance" 
is described 

0F=0F=2()1023sasaqKεF2()1023ssqKεFsas()qKε()qKεqεHHHH1ininˆH(,)h40 
 
 
 
 
   
 
 
 
 
 
 
 
 
where 
first approximation one may assume a constant value of 
plastic  strain.  The  parameter 
deformation. In the initial formulation of the Tseng – Lee model, 

is the hardening modulus of the memory surface when the two surfaces get in contact. As a 
can also be a function of equivalent 
affects  the  steepness  of  the  strain-strain  curve  during  plastic 

is assumed to be constant. 

, but 

The material model with the TL hardening rule has been implemented in a user-subroutine (UMAT) for 
ABAQUS,  using  an  Euler-forward  numerical  integration  scheme.  The  corresponding  material  model 
parameters have been properly defined based on the material testing data from CSM [13], presented in 

 have been selected equal to 5,000 MPa, 
another work package of the project. The values of 
and 10, to match the experimental results. In addition, an exponential function for controlling the size of 
the yield surface has been adopted: 

 and 

where 

  is  the  initial  yield  stress, 

and 

are  appropriately  defined  parameters.  In  the  present 

analysis, the initial size 

 of the yield surface is equal to 280 MPa, 

is -50 MPa (cyclic softening) 

 is equal to 100. The initial size of the memory surface is equal to 430 MPa, allowing for the 
and 
simulation of the smooth transition from the elastic to the plastic regime observed in material testing. 
The prediction of the stress-strain curve with the TL model is presented in Figure II.4.1-8. 

(7) 

Figure II.4.1-7: A schematic view of the two-surface 
model. 

Figure II.4.1-8: Material curve and model predictions 

II.4.2. 

II.4.2.1.

Task  4.2  Numerical  simulation  of  the  experimental  tests  on  components  of 
critical importance, tested in WP3 

Numerical  simulation  of  elbow  behaviour  under  extreme  monotonic  and  cyclic 
conditions 

A detailed finite element model has been developed to support the experimental results in the general-
purpose finite element code ABAQUS. Eight-node incompatible mode (C3D8I) solid elements are used 
for  the  simulations.  The  model  takes  into  account  geometrical  and  material  nonlinearities  through  a 
large-strain formulation and a von Misses description of the yield surface. The exact configuration of 
the two support systems at the specimen ends have also been modeled, to account for their flexibility in 
a rigorous manner. 

In Figure II.4.2-1 a general view of the numerical model is depicted. The curved part of the elbow and 
the two straight parts have been modeled as separate parts of the model, so that an appropriate value of 
thickness  is  assigned  at  each  part,  according  to  thickness  measurements.  For  each  test,  a  detailed 
numerical  model  has  been  developed,  adopting  the  mean  thickness  for  each  part  of  the  specimen 
(straight pipe segments and elbow) according to thickness measurements. The weld regions have been 
also modeled separately considering the local increase of thickness due to the cap of the weld.  

ˆˆHˆHhhˆHh0()1qbqkkQe0kQb0kQbmemory surfaceyield surfacennnssaorigin23k23k41 
 
  
 
 
 
 
 
For the description of the nonlinear elastic-plastic material behavior two advanced plasticity models, (a) 
the Armstrong-Frederick model and (b) the Tseng-Lee model, have been adopted, both based on von 
Mises plasticity.  

Figure II.4.2-1: General view of the numerical model 

Figure  II.4.2-2: Concentrations of plastic deformations 
– closing bending loads 

II.4.2.1.1. 
under strong cyclic loading 

Numerical  results  and  comparison  with  test  data  of  non-pressurized  piping  elbows 

Eight  models  have  been  developed  on  8-inch  SCH40  long-radius  induction  bends,  tabulated  with  the 
amplitude end-displacement and the number of cycles until failure in Table II.4.2-1 for each specimen. 

Simulation  results  for  Test  No4  are  compared  with  the  corresponding  experimental  measurements  in 
Figure  II.4.2-3.  The  comparison  between  the  experimental  results  and  the  numerical  predictions  is 
considered satisfactory for all the tests and is presented in Deliverable4. Using the finite element model 
and  adopting  the  most  appropriate  set  of  material  model  parameters  for  each  case,  the  load-
displacement  and  flattening-displacement  experimental  curves  can  be  reproduced  in  a  precise  and 
effective way.  In Figure II.4.2-4 the predictions of the numerical model concerning the evolution of the 
cross-section  flattening  in  “Position  1”  are  compared  with  the  experimental  measurements.  The 
numerical  predictions  are  in  good  agreement  with  the  measured  cross-sectional  flattening.  The 
concentration  of  plastic  deformations  reaches  its  highest  value  below  the  central  axis  of  the  elbow 
(Figure II.4.2-2).  

The evolutions of maximum and minimum local strains are of particular importance. It is presented in 
Figure II.4.2-5 and Figure II.4.2-6, based on the experimental and numerical results of the Test No.4. 
Since the behavior of the elbow is not symmetrical under opening and closing bending loads, the local 
mean strain is nonzero causing ratcheting phenomena. Due to these phenomena, despite the fact that the 
strain  range  remains  almost  constant,  there  is  an  increase  of  the  value  of  the  maximum  strains  and 
consequently an increase of the mean strain accelerating fatigue failure.  

The  numerical  predictions  for  the  evolution  of  local  strains  using  the  NLKH  and  the  TL  model  are 
compared  with  the  measured  values  and  are  found  to  be  in  reasonable  agreement  with  the 
experimentally  measured  values  (Table  II.4.2-1).  It  can  be  observed  that  the  NLKH  model 
overestimates the ratcheting  behavior  of the  material,  while the TL  model provides  strain  predictions 
much  closer  to  the  measured  values.  This  can  be  attributed  to  the  fact  that,  under  increasing  plastic 
deformations, the hardening modulus in the NLKH model constantly decreases resulting to higher strain 
predictions for a small increase of the stress level. 

42 
 
 
 
 
 
 
Table II.4.2-1: Displacement range, number of cycles and local strain range. 

Test 
No. 

Displ. range 

(mm) 

1 

2 

3 

4 

5 

6 

7 

8 

±25 

±70 

±100 

±150 

±200 

±250 

±300 

ECCS 
protocol 

Number of 
cycles until 
failure 

13160 

444 

171 

61 

28 

17 

10 

16 

Local strain 
range* (%) 

Local strain 
range* (%) 

Local strain 
range* (%) 

(Exp)  

(NLKH)  

(TL) 

0.33 

1.23 

- 

2.61 

- 

3.84 

4.02 

- 

0.46 

1.07 

1.59 

2.21 

2.66 

3.30 

4.29 

- 

0.33 

1.25 

1.59 

2.55 

2.77 

3.75 

4.03 

- 

*: calculated in the elbow hoop direction at the critical location 

Figure II.4.2-3:  Numerical simulation of Test 4 - Load-
displacement cycles 

Figure  II.4.2-4:    Numerical  simulation  of  Test  4  -  
Flattening- displacement cycles 

Figure  II.4.2-5:    Numerical  simulation  of  Test  4  - 
Evolution of maximum local strains 

Figure  II.4.2-6:  Numerical  simulation  of  Test  4  - 
Evolution of minimum local strains 

II.4.2.1.2. 
under strong cyclic loading 

Numerical  results  and  comparison  with  test  data  of  non-pressurized  piping  elbows 

Based on the properly calibrated numerical models presented in the previous sections, the tests reported 
in Table II.4.2-2 have been simulated. Indicative simulation results for tests 1 and 2 are compared with 
the  corresponding  experimental  measurements  in  Figure  II.4.2-7  and  Figure  II.4.2-9.  As  in  the 

lfN43 
 
 
 
 
 
 
 
 
experimental  set of  the  non-pressurized  elbows, the developed  numerical  models are  able  to  simulate 
accurately the tests on pressurized elbows as well. 

Table II.4.2-2: Displacement range and number of cycles 

Test 
No. 

Displ. range 

(mm) 

Pressure load [MPa] 

Number of cycles 
until failure 

1 

2 

3 

4 

5 

±200 

±200 

±200 

±300 

±300 

3.2 

7 

12 

3.2 

7 

26 

27 

22 

10 

10 

Figure  II.4.2-7:  Numerical  simulation  of  Test  No1-  
Load - displacement 

Figure  II.4.2-8:  Numerical  simulation  of  Test  No1  - 
Ovalization - displacement curve 

Figure II.4.2-9: Numerical simulation of Test No2: Load - 
displacement 

Figure  II.4.2-10:  Concentration  of  strains  at  the 
elbow critical region 

The  critical  position  where  cracking  took  place  in  the  experiments  reported  in  the  present  study  is 
verified by the numerical model as well. Bellow the elbow mid-axis at the elbow flank region, there is a 
localization of strains mainly in the hoop direction (Figure II.4.2-10). This justifies the fact that all the 
cracks formed along the longitudinal direction of the elbow. 

lfN44 
 
 
 
 
 
 
 
 
 
 
II.4.2.1.3. 

Fatigue life and fatigue curves 

Non-pressurized specimens Nos.2-7 have been subjected to constant amplitude end-displacement range 
and  failed  under  a  small  number  (less  than  105)  (see  Table  II.4.2-1).  Test  No.1  is  not  considered 
13,160) falls at the transition region between low-
because the corresponding number of cycles (

cycle and high-cycle fatigue. In addition, test No.8 refers to variable amplitude loading and, therefore, 
is not considered in this analysis. 

To examine the low-cycle fatigue behavior of the elbows under consideration, it is necessary to express 
  until  failure  in  terms  of  the  hoop  strain  amplitude  at  the  critical  location 
the  number  of  cycles 

where  cracking  takes  place.  The  numerical  results  have  been  in  good  agreement  with  the  available 
measured  strains.  For  each  test,  the  hoop  strain  range 
  at  the  critical  location  computed  from  the 
numerical simulation using either the NLKH model or the TL model is shown in Table II.4.2-1. 

Using these values for tests Nos. 2-7, the following fatigue equation can be derived that correlates the 
. According to the values reported 
number of cycles to failure 

 with the local hoop strain range 

in Table II.4.2-1, it is possible to derive equations result from the experimentally measured values and 
the corresponding values from the numerical analysis using the NLKH and the TL model respectively. 
The corresponding equations are:   

Experimental measured values 

Numerically predicted values - NLKH model 

Numerically predicted values - TL model 

       (4) 

     (5) 

     (6) 

The resulting fatigue curves are presented in Figure II.4.2-11. In this graph, the ASME BPVC standard 
[4]  which  provides  design  fatigue  curves  for  elbows  failing  in  the  low-cycle  fatigue  regime  is  also 
included.  

Figure II.4.2-11: Local hoop strain range versus number of cycles to failure. 

fNfNfN3.119.124fNε2.889.057fNε3.138.693fNε45 
 
 
 
 
 
 
 
 
II.4.2.1.4. 

Numerical simulation of pseudo-dynamic loading on piping system 

As  reported  in  Task  3.4,  a  pseudo-dynamic  test  campaign  was  carried  out  by  UTrento  in  order  to 
investigate  the  seismic  performance  of  a  typical  full-scale  industrial  piping  system  and  some  of  its 
critical  components.  In  a  Pseudo-Dynamic  test  with  Dynamic  Substructuring  (PDDS),  the  overall 
response of a structure is evaluated by combining the experimental response of a Physical Substructure 
(PS)  -which  is  generally  the  most  critical  part  of  the  structure-  with  the  numerical  response  of  a 
Numerical Substructure (NS) that contains the remainder of the structure. A typical full-scale industrial 
piping system, see Figure III.3.4-24, containing 8 and 6 inch straight pipes, several elbows, a Tee joint 
and a bolted flange joint was considered for the experimental tests. A 3D Finite Element (FE) model of 
the  piping  system  was  developed  in  SAP2000  [50]  software  in  order  to  perform  numerical  analyses. 
Two coupling nodes were chosen in the locations of bending moments close to zero found from a time-
history seismic analysis; see Figure III.3.4-24(b) for the two substructures and coupling nodes. It was 
shown  that  by  satisfying  the  compatibility  condition  only  in  the  horizontal  x  direction,  see  Figure 
II.4.2-12,  we  were  able  to  reproduce  the  response  of  the  original  system;  details  can  be  found  in 
Deliverable D3. The input earthquake accelerogram for analyses and experimental tests were generated 
on  the  elevated  floor  of  the  support  structure  of  the  piping  system  by  imposing  an  earthquake  to  its 
base.  PGAs  of  different  earthquake  levels  were  adjusted  according  to  the  serviceability  and  ultimate 
limit  states  suggested  by  an  Italian  Standard.  An  elastic  seismic  analysis  on  the  piping  system  was 
carried out in SAP2000 [50] with and without considering damping. The Rayleigh damping model was 
constructed  for  the  analysis  considering  a  0.5%  damping  for  the  piping  system;  it  was  estimated 
through identification tests on the test specimen. It was found that the Rayleigh’s damping considerably 
reduced the stress levels in the piping system.  

(a)  Specifications of the piping system 

(b)  PS, NS and coupling nodes 

Figure II.4.2-12: The piping system and the two substructures 

Implementation of the PDDS on the piping system was challenging mainly because it was a structure 
with  distributed  mass  –for  which  PDDS  has  been  considered  inadequate-  and  it  was  subjected  to 
distributed  earthquake  forces.  The  earthquake  forces  had  to  be  condensed  to  the  two  coupling  nodes 
owing  to  the  infeasibility  of  applying  earthquake  forces  in  each  node  of  the  PS  through  actuators. 
Additionally, mass and damping  matrices of the PS were also to be reduced to 2 by 2 matrices to be 
numerically modelled in the system of equations of motion of the piping system. The so called Craig-
Bampton [18] mode synthesis technique -which was found to be the most effective among others- was 
adopted  to  accomplish  this  task.  The  LSRT2  integration  algorithm  developed  by  Bursi  et  al.  [9]  was 
used to perform the PDDS. Tests were carried out by using two MOOG hydraulic actuators, an MTS 
FT60  controller  [44],  an  xPC  target  and  a  Host  PC.  The  hardware-software  scheme  for  the  PDDS  is 
presented in Figure II.4.2-13. 

46 
 
 
 
 
 
 
Figure II.4.2-13: Hardware-Software configuration for the PDDS 

Several  PDDS  were  carried  out  on  the  piping  system  under  different  levels  of  earthquake  PGAs 
corresponding  to  serviceability  limit  states  suggested  by  Standards;  refer  to  Task  3.4  for  the 
experimental program and input earthquake PGAs. All experimental results showed a good agreement 
with  relevant  simulation  results  as  can  be  noted  from  Figure  II.4.2-14  in  which  experimental  and 
numerical  displacements  in  Coupling  Node  #1  under earthquake  PGAs  corresponding  to  the  safe  life 
limit state (SLVT) and collapse limit state (SLCT) are compared. 

(a)  

(b) 

Figure II.4.2-14: Displacement histories and relevant spectra of the Coupling Node #1 at (a) SLVT and (b) SLCT 
(Blue: Experimental; Magenta: Numerical) 

II.4.2.2.

Numerical simulation of tank nozzles  

Subject of these simulations was numerical recalculation and verification of 6 tests under strong cyclic 
loading, as carried out and described in Task 3.2. The geometries of the FE-models were chosen equal 
to  the  measured  dimensions  of  the  test  specimens  (including  tolerance  deviations).  Figure  II.4.2-15 
shows an overview of the FE-models, whereby the numerical simulations were conducted taking into 
account advantages of symmetry conditions. 

47 
 
 
 
 
 
 
 
(P1) 

(P2) 

(P3) 

Figure II.4.2-15: FE-models overview: (P1) internal-, (P2) external- and (P3) nozzle-reinforcement. 

The  FE-program  ABAQUS  (version  6.12.1)  was  used  and  the  FE-models  were  created  using  8-node 
reduced-integration  solid  elements  (type  C3D8R).  Both  geometric  and  physical  non-linearities  were 
considered  and  moreover  the  FE-mesh  as  well  as  the  FE-mesh  refinements  around  the  critical  points 
were calibrated to get stable and adequate results. 

To  have  accurate  input  data  for  the  material  characteristics,  some  small  scale  specimens  were  taken 
from the material of the shell plates of the nozzle specimens (P355NH). These small scale specimens 
were tested at CSM under monotonic tensile (4 tests) and cyclic loading (3 tests).  For the large scale 
simulations  under  monotonic  loading  conditions,  in  order  to  determine  the  test  procedure  for  cyclic 
loading according to ECCS [20] (see Task 3.2), non-linear isotropic hardening was considered using the 
extrapolated true stress-strain curve obtained from the test results of the small scale tests, whereas for 
cyclic  loading  conditions  linear  isotropic  and  non-linear  kinematic  hardening  according  to  Lemaitre 
Chaboche  [42]  were  considered.  The  required  parameters  were  determined  analytically  and  were 
verified numerically using numerical simulations of the small scale tests under cyclic loading. 

With regard to the test setup the  lateral supports as well as the load application were incorporated by 
adequate boundary conditions. Also the possibility of the bottom plates to lift up was allowed for by 
using a “surface to surface” contact definition between the bottom plates and support.  The numerical 
simulations were run displacement-controlled with the measured loading history of the tests. 

The numerical verification of the tests was aiming at two main aspects: 

  Validation of the numerical model with regard to its accordance with the experimental 

behaviour and thus its applicability for parametric studies; 

  Identification of damage and formulation of equivalent damage criterion transferrable from 

damages observed in the tests to the numerical model. 

The experimental and numerical load-displacement curves show a very good agreement confirming the 
capability of the FE-models to represent the overall performance of the shell nozzle reinforcements. For 
the  identification  of  failures  (cracks)  in  the  numerical  FE-models  an  approach  has  been  selected,  in 
which the accumulated plastic strains ε p were compared to a limit value ε p,limit. The limiting values have 
been  calibrated  by  the  test  results  such  that  for  each  tested  specimen  respectively  its  FE-model  the 
location  and  the  accumulated  plastic  strains  have  been  determined  from  the  simulations  at  the  cyclic 
deformation stage of the tested specimen corresponding to the first visually detected crack. From this 
evaluation the number of cycles until failure Nf and the location (point of failure) Pf can be obtained. 

The limiting plastic strain is not universally applicable. By applying this method two different values 
have been obtained – ε p,limit = 8.1 % for longitudinal and ε p,limit = 2.2 % for transverse load direction. For 
each load direction however one value matched all three types of shell nozzle reinforcement. 

The experimental and numerical results for Nf and Pf are summarized in Table II.4.2-3. 

48 
 
 
 
 
 
Table II.4.2-3: Experimental and numerical results for Nf and Pf 
Nf,num 
[-] 
23.1 

Nf,exp 
[-] 
24.0 

Reinforcement 

P1-1 Internal 

P2-1 External 

P3-1 Nozzle 

P1-2 Internal 

P2-2 External 

24.0 

16.0 

9.0 

9.0 

24.4 

19.5 

8.4 

8.7 

1) 

Pf,exp
[-] 
1,2 

1,2 

3 

1 

1 

1) 

Pf,num
[-] 
1 

1 

3 

1 

1 

P3-2 Nozzle 
1)  Point  of  failure:  1  -  nozzle;  2  -  reinforcing  plate;  3  -  shell 
plate 

11.0 

10.1 

3 

3 

II.4.2.3.

Numerical simulation of flanged joints 

II.4.2.3.1 

Numerical simulation of standard flanged joints 

In this section are reported the results of the FE numerical simulations of the tests on standard EN1092-
1  [23],  bolted  weld  neck  flanged  joints.  The  aim  of  these  simulations  is  to  obtain  a  model  that 
reproduces with accuracy the behavior of the specimen subjected to monotonic and cyclic bending with 
internal pressure.  
A FE model has been developed by adopting the commercial software MSC.Marc©. Some details of the 
mesh are reported in Figure II.4.2-16. In order to obtain a good balance between the accuracy required 
and the computational resources request, it was thus decided, exploiting the symmetry, to model only ¼ 
of  the  entire  geometry,  and  to  adopt  different  element  types  for  different  regions  of  the  model.  The 
flange has been modeled with 3D brick elements, the bolt shank by means of beam elements and the 
bolt head is modeled as a surface that goes in contact with the  flange ring. Solid-shell elements have 
been chosen  for the  mesh  of  the  pipe  body This  type  of  element  uses  the  8 node  brick  topology,  but 
adopts  a  shell  element  formulation  which  is  particularly  adequate  to  represent  the  behavior  of  pipes 
subjected  to  bending  in  combination  with  internal  pressure  and  provides  higher  computational 
efficiency if compared to 3D brick elements. The gasket has been modeled in detail by means of 3D 
brick  elements.  In  particular  the  spiral  wound  sealing  element  adopts  a  “gasket”  material  model 
provided in Marc, which allows the use as input of pressure vs. closure relationships which describe the 
trough thickness behavior of the gasket. 

Figure II.4.2-16: Mesh details 

The P355 N material true stress true strain curve obtained by full thickness specimens extracted from 
pipe  in  longitudinal  direction  have  been  employed  for  the  simulation  of  the  monotonic  tests  and  is 
reported  in  Figure  II.4.2-17a.  For  the  simulation  of  the  cyclic  tests,  nonlinear  kinematic  hardening 
model  have  been  adopted. The  material  uniaxial  behaviour  introduced  as input  in  FEA  is  reported in 

49 
 
 
 
 
 
 
Figure II.4.2-17b. This model can better represent the actual material behaviour when cyclic loading is 
present as it is capable, even with some limitations, to reproduce the Bauschinger effect, which on the 
contrary is totally neglected in isotropic or linear kinematic hardening material descriptions. 

Figure II.4.2-17: Monotonic and cyclic material properties employed in FEA 

(a) 

(b) 

Two typical  numerical  vs. experimental  results comparison  are reported  in  Figure  II.4.2-18  for  PN40 
monotonic  bending  test,  and  Figure  II.4.2-19  for  the  PN63  connection  subjected  to  a  bending 
displacement of +/-27mm. It is possible to observe that the model is able to capture the bending with 
very good accuracy all phases of the monotonic case: yielding, plastic behavior, and local buckling of 
the pipe which causes a load drop due to collapse of the section. In case of the cyclic bending loading a 
good agreement is obtained although the material model adopted in FEA provides a sharper transition 
between  elastic  to  plastic  regime  with  respect  to  smooth  and  continuous  transition  experimentally 
observed. 

Figure II.4.2-18: PN63-1 FE model deformed shape and plastic strains at maximum bending deflection (right) and 
specimen at the end of the test (left) 

Figure  II.4.2-19:  PN63-1  FE  model  deformed  shape  and  plastic  strains  at  maximum  bending  deflection  (right) 
and specimen at the end of the test (left) 

01002003004005006007000.000.040.080.120.16True Stress [MPa]True strain [-]-600-400-2000200400600-0.03-0.02-0.02-0.01-0.010.000.010.010.020.020.03True Stress [MPa]True strain [-]0501001502002500100200300400500050100150Deflection (mm) Load [kN]Stroke [mm]Experimental loadFEA LoadExperimental DeflectionFEA Deflection-600-400-2000200400600-30-20-100102030Load [kN]Stroke [mm]FEAExperimental50 
 
 
 
 
 
 
 
 
 
 
II.4.2.3.2 

Numerical simulation of non-standard flanged joints 

With reference to bolted flanged joints, the aim was to calibrate an accurate numerical model of non-
standard  flanged  joints  on  the  basis  of  experimental  results.  In  order  to  obtain  these  results  a  finite-
element software ABAQUS 6.11-2 ([1], [2]) has been used. In fact, this software allowed to build a 3D 
model of pipes, flanges and bolts by means of an eight-node hexahedron brick elements called C3D8, 
and  allowed  to  understand  the  behaviour  of  joints  under  monotonic  loads.  UTrento  designed  two 
different non-standard joints for testing. The thicknesses were chosen according to the failure modes 1 
and 2 of the Eurocode EN 1993-1-8 [29]. The thicknesses of the designed non-standard flanges were 
taken as 18 mm (Design 01) and 27 mm (Design 02), respectively. Apart from the thicknesses, all other 
dimensions of the designed flanges were the same as those of a PN 40 (for a DN 200 pipe size) plate 
flange given in the Eurocode EN 1092-1 [23]. 

The actual joint system was modelled with the ABAQUS software, as shown in  Figure  II.4.2-20. For 
bending tests a fixed-roller end constraints scheme was adopted and the load was applied by means of 
two plates that surrounded pipes; instead, for axial test, a free-fixed end constraints scheme was used, 
while the load was applied on the free end through the upper steel plate. 

(a) 

(b) 

Figure II.4.2-20: Geometrical models for the: (a) bending test, and (b) axial test 

Meshing was obtained through C3D8R  - with reduced integration - and C3D8I  - with full integration 
and  incompatible  modes-  brick  elements.  The  non-linearity  contacts  were  shaped  by  the  so-called 
Interactions that identified a Master surface and a Slave; this wording recognizes that the surfaces were 
in  contact,  requiring  that  the  nodes  on  the  slave  surface  could  not  penetrate  each  other  in  the  master 
surface.  The  sliding  and  the  contact  responded  to  the  Coulomb’s  friction  law.  The  welds,  which 
guaranteed the connection between the flange and the tube, were shaped by a Tie contact, as well as the 
contact  between  the  stem  of  the  bolt  and  nuts  and  between  these  and  the  flange.  The  analysis  of  the 
responses of the Bending Test Design 01 and 02 – see Figure II.4.2-21(a) and Figure II.4.2-21(b) – and 
of  the  Axial  Test  Design  02  –  see  Figure  II.4.2-21(c)  –  shows  that  responses  obtained  with  the  FE 
model  appear  to  be  in  agreement  with  experimental  curves.  However,  for  the  Bending  test,  the 
behaviour in the elastic range was not completely caught by the model; this phenomenon is mainly due 
to the fact that the boundary conditions of the experiments are difficult to be approximated in the FE 
models.  For  the  Axial test,  instead, the behaviour  of the  numerical  model in the  elastic range is  very 
close to that experimental one. Beyond the elastic limit the two curves diverge, and bricks with reduced 
integration elements provide a response less strong than the elements with incompatible models. 

The gasket between flanges has not been taken into account in the model. The reason is because this 
element  does  not  significantly  influence  the  overall  response  of  the  FE  models,  though  determines 
leakage. The difference between the experimental curve and the responses of the numerical models is 
also linked to the fact that all various imperfections due to working or the alteration of properties of the 
materials, like residual stresses, were not considered in the numerical analysis. 

51 
 
 
 
 
 
(a) 

(b) 

(c) 

Figure  II.4.2-21:  Comparison  between  experimental  and  numerical  data  for  bending  tests:  (a)  Design  01,  (b) 
Design 02; and axial tests: (c) Design 01 

A  comparison  between  the  model  that  contained  C3D8R  elements  and  the  one  with  C3D8I  elements 
was done. The first ones are the preferable choice for the bending model, while the latter are the best 
choice for the axial model, according to results shown in Figure II.4.2-21.  

(a) 

(b) 

(c) 

Figure II.4.2-22: Deformed shapes and stress levels in the joint for the: (a) bending test – Design 01 –, (b) bending 
test – Design 02 –, and (c) axial test – Design 01 –. 

Finally, relationships are used to calculate the Normal Root Mean Square Deviation (NRMSD) and the 
Root  Mean  Square  Deviation  (RMSD)  of  the  error  between  experimental  results  and  numerical 
simulations. They read: 

where Fex is the force applied in the experimental case, Fnum the force provided by the numerical model 
and  N  the  number  of  load  steps.  Table  II.4.2-4  summarises  all  relevant  results  which  appear  to  be 
satisfactory. 

Table II.4.2-4: Error between experimental results and numerical simulations calculated with NRMSD and RMSD 
methods 

Test 

Bending 

Bending 

Axial 

Model 

Design 01 

Design 02 

Design 01 

NRMSD (-) 

RMSD (kN) 

0.13 

0.12 

0.10 

29.91 

25.94 

132.42 

2221;exnumexnumnumFFNRMSDRMSDFFNF52 
 
 
 
 
 
 
 
 
 
 
 
 
 
II.4.2.4.

Numerical simulation of Tee pipe joints 

The  experimental  investigation  of  the  junction  performance  (WP3)  is  supported  by  detailed  finite 
element  models,  developed  to  simulate  the  experiments.  Two  finite  element  models  have  been 
developed  in  ABAQUS  FE  code  for  each  specimen  type  using  8-noded  reduced  integration  solid 
elements  (C3D8R).  The  symmetry  of  the  problem  was  taken  into  consideration  allowing  for  the 
simulation of the half specimen. At the symmetry plane, the appropriate symmetry boundary conditions 
were applied. In addition, the extension part attached to the branch end facilitating the connection of the 
specimens  to  the  loading  device  was  also  modeled  accurately  in  order  to  take  into  consideration  its 
flexibility. Finally, a rigid beam (B31) element is attached at the end of the tubular member by a hinge 
interaction to simulate the connection between the hydraulic cylinder to the specimens. The loading is 
applied at the free end of the rigid beam which is allowed to be translated only in the vertical direction 
to  the  tee-junction  main  plane.  In  the  tubular  members  of  the  junction  three  elements  along  the 
thickness where used for the main pipes connected and four elements for the fitting component, while 
the  mesh  density  varies  along  the  length  of  the  member  as  presented  in  Figure  II.4.2-23  for  the  Tee 
branch  junction  with  the  fitting  and  the  reinforced  branch  junction.  The  nonlinear  behavior  of  the 
material  is  described adopting  the  advanced  cyclic  plasticity  material  constitutive  model  proposed  by 
Tseng-Lee [53]. 

The developed numerical model is used to simulate the tests reported in Table II.4.2-5.  The comparison 
of the  predicted load-displacement  curves  with  the corresponding  experimental  curves  shows  that  the 
adopted model is able to provide numerical results in good agreement with the experimental findings. 
Indicative results are presented in Figure II.4.2-24 and Figure II.4.2-26. The observed differences can 
be attributed to minor slips at the end supports of the test specimens. These slips are more evident in the 
TJF  tests  due  to  their  increased  load  resistance.  Nevertheless,  the  numerical  models  can  predict  the 
maximum and minimum resistance loads fairly well.  

In  addition,  the  stress  and  strain  concentration  areas  according  were  examined  using  the  developed 
numerical model. As it is presented in Figure II.4.2-25 in the TJF specimens there is a concentration of 
strains  at  the  curved  part  of  the  fitting  connecting  the  main  pipe  to  the  branch.  Due  to  the  local 
curvature,  the  distribution  of  the  developing  strains  is  spread  in  a  bandwidth  of  about  2  cm.  On  the 
contrary, in the TJWR specimens due to the existence of the welds, the strain concentration takes place 
at  the  weld  area  of  the  branch  pipe  (Figure  II.4.2-27).  The  above  observations  are  in  excellent 
accordance with the experimental findings. 

Specimen type 

Loading 
type 

Displacement 

(kN) 

  (mm) 

TJF 

TJWR 

1M 
2C 
3C 
4C 
1M 
2C 
3C 
4C 

M* 
C* 
C 
C 
M 
C 
C 
C 

200a 
60 
45 

158.0 
131.2 
117.9 
159.9  ECCS No45 
81.9 
78.3 
64.9 
77.9 

200a 
60 
45 
ECCS No45 

Cycles 
to 
failure 

- 
42 
135 
15 
- 
6 
17 
11 

*: M=monotonic, C=cyclic,  
a: Maximum stroke displacement reached 

Figure  II.4.2-23:  Numerical  models:  General 
view 

Table II.4.2-5: Experiments results 

All  the  Tee-junction  specimens  failed  due  to  a  through-thickness  crack.  The  TJWR  specimens  failed 
along the circumference of the branch-to-reinforcement plate weld-toe. The specimens examined under 
monotonic loading and under an imposed LPD of 60 mm, failed in the branch side of the connection 

maxPδfN53 
 
 
 
 
  
 
 
while the other two specimens in the welded reinforcement plate side of the connection. On the other 
hand, the TJF specimens failed in the region of the joint fitting, at a distance of about 60 to 80 mm away 
from the branch-to-fitting weld-toe.  

The developed numerical models were able to simulate fairly well the behavior of both types of Tee-
junctions. The areas where cracking took place are also indicated in the models. While in the case of 
TJWR the critical areas are located at the welds of the Tee-junction, in the case of the TJF type, there is 
significant stress and strain concentration at an area located about 80 mm from the weld region of the 
fitting component to the secondary tube branch.  

Figure  II.4.2-24:  Finite  element  analysis  results:  
TJF 2C test 

Figure II.4.2-25: Numerical TJF model  

Figure II.4.2-26: Finite element analysis results:  TJWR 
2C test 

Figure II.4.2-27: Numerical TJWR model   

The  current  numerical  analysis  focused  on  the  strength  and  fatigue  resistance  behavior  of  two 
configurations of Tee-junctions, one using a fitting component and a second one using reinforcement 
plates at the connection of its two main parts. Under monotonic loading conditions, the specimen with 
the Tee fitting showed about 93% higher load capacity at an imposed displacement of 200 mm. Under 
cyclic loading conditions, for the imposed displacement value of 60 mm, the TJF specimen exhibited 
higher fatigue resistance of about 7 times that of the TJWR one. For the imposed cyclic displacement of 
45 mm, the TJF specimen exhibited a higher fatigue resistance of about 8 times that of the TJWR one. 
For the low-cycle fatigue tests according to ECCS No45, the TJF specimen resisted 36% more loading 
cycles than the TJWR one. 

54 
 
 
 
 
 
 
 
 
II.4.3. 

Task 4.3 Parametric numerical study of critical component behaviour 

II.4.3.1.

Parametric  numerical  study  of  elbow  behavior  under  extreme  monotonic  and  cyclic 
conditions 

The  behaviour  of  8-inch  elbows  under  in-plane  and  out-of  plane  bending  is  investigated  through  a 
parametric study. The complete list of their geometrical characteristics is reported in  Table II.4.3-1. A 
detailed finite element model has been developed to support the above experimental results and conduct 
the  parametric  study  in  the  general-purpose  finite  element  code  ABAQUS.  Both  eight-node 
incompatible  mode  (C3D8I)  solid  elements  and  four-node  reduced  integration (S4R)  are  used for the 
simulations.  Both  types  of  elements  have  shown  a  very  good  performance  in  simulating  the 
experiments.  The  model  takes  into  account  geometrical  and  material  nonlinearities  through  a  large-
strain  formulation  and  a  von  Mises  yield  surface.  For  the  description  of  the  nonlinear  elastic-plastic 
material  behavior  two  advanced  plasticity  models  have  been  adopted,  both  based  on  von  Mises 
plasticity. These advanced models have been selected among other more simplified models (e.g. linear 
kinematic hardening model) because of their ability to describe more accurately the Bauschinger effect, 
which  is  evident  upon  load  reversal  of  the  material  stress-strain  curve  in  the  plastic  range  and  may 
affect  the  overall  behavior  of  the  elbow  specimens.  In  this  parametric  study,  the  two  pipe  ends  are 
capped,  and  a  bending  moment  is  applied.  The  aim  of  this  study  is  to  examine  the  effects  of  the 
variation of the elbow geometrical characteristics on its mechanical behavior. 

Table II.4.3-1: Geometrical characteristics of examined cases 

Geometrical 
Characteristics 

Diameter 
(mm) 

Thickness 
(mm) 

Bend radius
(mm) 

8-inch SCH80 

Long radius 

8-inch SCH80 

Short radius 

8-inch SCH40 

Long radius 

8-inch SCH40 

Short radius 

8-inch SCH20 

Long radius 

219.1 

12.7 

304.8 

1.391 

219.1 

12.7 

203.2 

0.927 

219.1 

219.1 

219.1 

8.2 

8.2 

6.4 

304.8 

1.391 

203.2 

0.927 

304.8 

1.391 

II.4.3.1.1. 

In-plane bending 

The pipe elbows have been subjected to monotonic in-plane bending first and subsequently to severe 
cyclic  in-plane  bending  through  the  application  of  cyclic  end-displacements  on  the  moving  support, 
keeping the other support pin-connected to the supporting steel frame. A general view of the numerical 
model is presented in Figure II.4.3-1. The specimens followed a constant amplitude end-rotation equal 
to ±0.02, ±0.04, ±0.08 and ±0.12 rad. 

In the monotonic loading cases, the elbow is subjected to opening or closing moment loading (Figure 
II.4.3-2). In this case, the two plasticity models predict similar behavior. Nevertheless, some differences 
between two models become evident at large rotation levels, where significant plastic deformations are 
involved, due to the different constitutive equations formulation of the two models. The same trend in 
the predictions of the two models is also observed in the cyclic loading at different end-rotation levels 
presented in Appendix III.4.1.1. Despite the fact that the resulting moment-rotation curves by the two 
models are close to each other, the measured local strain ranges differ significantly. Consequently, the 
predicted  fatigue  life  for  these  cases  may  differ.  The  reason  for  the  above  observations  is  that  under 

DtRR/D55 
 
 
 
 
increased  plastic  deformations,  the  hardening  modulus  predicted  by  the  NLKH  model  tends  to  zero, 
while the TL model provides a more realistic prediction for it.  

Figure II.4.3-1: Numerical model for the elbows used 
for the parametric analysis 

Figure  II.4.3-2:Numerical  parametric  analysis 
results  –  8-inch  SCH40  long  radius  -  Monotonic 
analysis 

II.4.3.1.2.  Out-of-plane bending 

The specimens have been also subjected to out-of-plane bending through the application of out-of-plane 
end-rotations  on  the  free  section,  keeping  the  other  section  fully  restrained.  A  general  view  of  the 
developed numerical model is presented in Figure II.4.3-3. The monotonic loading response is reported 
in  Figure  II.4.3-5  for  the  case  of  8-inch  SCH40  long  radius.  The  elbow  is  subjected  to  out-of-plane 
moment loading with three levels of pressure (0MPa, 3.2 MPa and 7.0 MPa). It should be noted that 
under  out-of-plane  bending,  the  elbow  is  subjected  to  a  combination  of  bending  and  torsion.  The 
presence  of torsion  is  responsible for  the  direction  of  the buckle,  whereas the additional compressive 
stresses  at  the  “flank”  of  the  elbow  due  to  bending  result  in  excessive  strain  development  at  this 
location.  Apparently,  the  buckle  occurs  at  the  “flattened”  area  of  the  cross-section,  where  buckling 
resistance is minimal due to the small value of local curvature (Figure II.4.3-4). For pressure equal to 
10% of py the buckled shape is similar to the one of the non-pressurized case. On the other hand, in the 
case  of  pressure  level  equal  to  20%  of  py,  in  cases  of  SCH80  and  SCH20,  the  elbow  distortion  is 
minimized, and buckle may appear at the straight part of the pipe. Based on those monotonic loading 
results, the specimens were subjected to out-of-plane cyclic following a constant amplitude end-rotation 
equal to ±0.05, ±0.1, ±0.25 and ±0.4 rad.  

Figure II.4.3-3: General view of the numerical model 

Figure II.4.3-4: Concentrations of plastic deformations 

56 
 
 
 
 
 
 
 
 
 
Figure II.4.3-5:Moment versus end-rotation curve for 
pressure 0, 3.2 and 7.0 MPa 

Figure  II.4.3-6:  Moment  versus  end-rotation  curve 
for displacement ranges ±0.05, ±0.1, ±0.25 and ±0.4 
rad, (pressure 0 MPa). 

II.4.3.2.

Parametric numerical study of tank nozzles 

Within this task the findings from the tests as well as from the simulations of the tests were broadened 
by further parametric numerical study. On the basis of the verified FE-models of the 3 investigated shell 
nozzle  reinforcements  the  influence  of  component  parameter  on  the  cyclic  behaviour  of  the  different 
shell  nozzle  reinforcements  was  assessed.  The  reference  geometries  have  been  represented  by  the 
nominal dimensions of the specimens. While all dimensions were kept unchanged, only one parameter 
was varied. The defined parameters are shown in Figure II.4.3-7 and summarized in Table II.4.3-2. 

Figure II.4.3-7: Overview of the geometrical parameters for the parametric study 

Table II.4.3-2: Geometrical parameters for the parametric study 

In the first step monotonic simulations for the 4 different load directions (see Task 3.2) for each type of 
reinforcement  and  each  variation  of  parameter  were  conducted,  thus  the  total  number  of  monotonic 
simulations resulted in 240 (+ 12 for the reference models). In order to assess the influence of internal 

12Ref34Shell thickness ts1015202530Bottom plate thickness tb58101520Nozzle thickness tn8.8101112.514.2Reinf. plate thickness tr58101520Reinf. plate diameter dr250275300350400as above ts , tb , tnReinf. nozzle thickness tnr17.52022.22528Reinf. nozzle length lr5062.575100150ParameterVariations of parametersP1 Internal and P2 External [mm]P3 Nozzle [mm]similar to P1 and P257 
 
 
 
 
 
 
 
 
 
hydrostatic pressure, the same monotonic simulations were carried out under consideration of internal 
hydrostatic  pressure  (2  bar  corresponding  to  a  tank  height  of  20  m).  By  comparing  all  parametric 
numerical  results for  monotonic  loading  it  could  be  observed  that the  differences  between  the  results 
obtained  with  and  without  consideration  of  internal  hydrostatic  pressure  are  small  but  also  remain 
nearly  constant  for  each  varied  parameter.  This  observation  led  to  the  conclusion  that  the  cyclic 
simulations  without  considering  of  internal  hydrostatic  pressure  would  provide  sufficient  information 
on the dependency of the performance as function of the varied parameter. 

In  the  second  step  cyclic  simulations  for  longitudinal  and  transverse  load  direction  and  each  type  of 
reinforcement  were  conducted.  Contrary  to  the  monotonic  simulations,  here  only  simulations  for 
parameter  variation  1  and  4  (+ reference)  according  to  Table  II.4.3-2  were  considered,  thus  the  total 
number of cyclic simulations resulted in 60 (+ 6 for the reference). 

The numbers of cycles until failure Nf  and the corresponding amplitudes of displacement emax as well as 
the  observed  points  of  failure  Pf    determined  as  described  in  Task  4.2.3  are  summarized  in  Table 
II.4.3-3 for longitudinal load direction and in Table II.4.3-4 for transverse load direction. 

Table II.4.3-3: Parametric numerical results for cyclic loading in longitudinal load direction 

Table II.4.3-4: Parametric numerical results for cyclic loading in transverse load direction 

NfemaxPf1)NfemaxPf1)NfemaxPf1)[-][mm][-][-][mm][-][-][mm][-]Ref.20.970.0122.680.0118.560.03ts = 1024.580.0228.3100.02,318.460.03ts = 3014.350.0115.650.0118.760.03tb = 521.370.0123.280.0118.160.03tb = 2020.070.0120.870.0119.670.03tn = 8.818.160.0119.270.0118.460.03tn = 14.222.380.0223.880.02,318.660.03tr = 52)21.470.0223.980.0218.660.03tr = 202)19.870.0121.270.0118.560.03dr = 2503)21.170.0122.880.0118.460.03dr = 4003)20.270.0122.480.0119.170.03Parameter [mm]P2-1 ExternalP3-1 NozzleP1-1 Internal1) Point of Failure: 1 - nozzle; 2 - reinforcing plate; 3 - shell plate2) P3-1: tnr = 17.5 / 28 mm instead of tr = 5 / 20 mm3) P3-1: lr = 50 / 150 mm instead of dr = 250 / 400 mmNfemaxPf1)NfemaxPf1)NfemaxPf1)[-][mm][-][-][mm][-][-][mm][-]Ref.14.450.0115.150.0117.660.03ts = 1022.180.0222.680.02,319.270.03ts = 3011.240.0110.740.0117.060.03tb = 515.650.0116.160.0117.560.03tb = 2013.550.0114.150.0118.260.03tn = 8.813.150.0113.750.0117.460.03tn = 14.217.260.0119.170.0117.960.03tr = 52)16.060.0116.160.0117.760.03tr = 202)13.750.0114.250.0117.660.03dr = 2503)14.650.0115.550.0117.460.03dr = 4003)14.250.0114.750.0117.860.031) Point of Failure: 1 - nozzle; 2 - reinforcing plate; 3 - shell plate2) P3-1: tnr = 17.5 / 28 mm instead of tr = 5 / 20 mm3) P3-1: lr = 50 / 150 mm instead of dr = 250 / 400 mmParmeterP1-2 InternalP2-2 ExternalP3-2 Nozzle58 
 
 
 
The  observations  and  conclusions  given  hereafter  summarize  the  results  of  the  numerical  parametric 
study  but  also  consider  the  results  of  the  experimental  tests  and  monotonic  simulations.  The  first 
observation regards the experimental tests on the 3 types of reinforcements. Despite differences in detail 
regarding the load-deformation characteristics and the localisation of failure (crack formation), all types 
showed  a  clearly  ductile  behaviour  and  resulted  to  remarkably  high  number  of  cycles  at  high 
displacement amplitudes. 

Fundamental informations on the behaviour characteristics are available already from monotonic load-
displacement  curves.  They  provide  information  on  stiffness  and  yield  load.  Under  the  condition  of 
displacement-controlled  cyclic  loading  increased  stiffness  leads  to  lower  cyclic  performance  and 
respectively  lower  number  of  cycles  until  failure;  the  cycles  however  are  coupled  with  higher  load 
amplitudes. 

Summarizing the performance of the 3 types of reinforcement observed from the tests and simulations 
the following conclusions can be drawn: 

  Type 1  and  2  (internal  and  external  reinforcement)  provide  a  comparable  performance  under 

cyclic loading; 

  Type 1 and 2 show a favourable performance under longitudinal loading; 
  Type 1  and  2  exhibit  the  best  cyclic  performance  in  combination  with lower  thickness  of the 

tank shell and of the bottom plate; this fact makes them a preferable solution for smaller tanks; 
  The detailing requirements for type 1 and 2 provided by EN 14015 [26] leave some space for 
the  selection  and  combination  of  parameters  within  defined  limits;  larger  diameter  dr  and 
smaller thickness tr of the reinforcing plate leads in general to higher cyclic performance. The 
location of the damage however moves from the nozzle (thick reinforcement) to the reinforcing 
plate (thin reinforcement); this can be of disadvantage for the stability of the tank shell and for 
the reparability of the damage. 

  Type 3 (thickened nozzle) show a favourable performance under transverse loading; 
  Type 3 perform in particular better than type 1 and 2 when a thick tank shell and bottom plate is 

present; this fact makes it more favourable for large tanks; 

  The possible increasing of thickness tnr and length  lr of the thickened nozzle of type 3 has no 
significant  influence  on  the  cyclic  performance;  for  each  investigated  parameter  however  the 
damage was always observed in the shell plate of the tank; 

  Regarding the stability of the tank shell a solution with a larger nozzle thickness tnr and shorter 

nozzle length lr in combination with a significantly thinner attached pipe is preferable. 

The provisions of EN 14015 [26] for the reinforcement of shell nozzles -which are intended to be used 
for static design - can be applied to shell nozzles under seismic actions as well. Additional measures, 
e.g.  increase  of  reinforcing  thicknesses,  does  not  necessarily  increase  the  performance;  quite  the 
contrary the performance can decrease. 

II.4.3.3.

Parametric numerical study of flanged joints 

II.4.3.3.1 

Parametric numerical study of standard flanged joints 

A FEA parametric study has been carried out on EN1092-1 [23] type 11 welding neck standard flanged 
joints  subjected  to  bending  loading.  The  objective  of  this  study  is  to  investigate  the  effect  of  the 
parameters  such  as  internal  pressure  level,  axial  load  and  geometry  on  the  behaviour  of  the  bolted 
flanged  connections.  Three  geometries,  as  reported  in  Table  II.4.3-5,  have  been  taken  into  account, 
while the load cases considered are reported in Table II.4.3-6. A schematic of the load application on 
FE mode is depicted in Figure II.4.3-8. 

59 
 
 
 
 
Table II.4.3-5: Flanged joints considered in parametric study 

Joint type 

Geometry 

ODxWT [mm] 

Pipe size 

Pipe D/t 

Steel grade 

- 

35 

27 

17 

- 

P355N 

P355N 

P355N 

EN1092-1 Standard  
weld neck  type 11 

DN200 - PN40 

219.1 x 6.35 

DN200 - PN63 

219.1 x 8.18 

DN200 - PN100 

219.1 x 12.7 

P=0.5*PN, N=+0.12*Ny 

P=0.5*PN, N=+0.24*Ny 

P=0, 
N=0 

P=0.5*PN, N=0 

P=PN, 
N=0 

P=0.5*PN, N=-0.24*Ny 

P=0.5*PN, N=-0.12*Ny 

notes 

Ny = pipe section yield axial load 

PN =maximum allowable pressure 

Table  II.4.3-6:  Load  cases  analyzed  for  each 
geometry 

Figure II.4.3-8: FE model and load scheme considered in 
parametric study 

The  parametric  study  confirmed  that  for  standard  weld  neck  flanged  joints  subjected  to  bending  the 
resistance of the pipe and flange assembly is governed by the pipe characteristics. On the flange, the 
only large strain spots can be reported on the neck, close to the weld, where neck thickness approaches 
the pipe thickness. On the pipe some large strain spots can be found on the extrados (tensile strains) and 
on the bulged sections on the intrados (compressive strains). 

Internal pressure, the axial load and the geometry have  marked effect in the joint bending resistance, 
which  has  been  identified  by  the  maximum  of  the  bending  moment  and  the  corresponding  rotation. 
Internal pressure has a beneficial effect on the sustainable rotations before instability of the joint, while 
its effect on the maximum bending moment is always negligible, up to the maximum operative pressure 
allowable by EN1092-2-1 code. On the contrary local tensile strains on pipe extrados and flange neck 
tend to increase with internal pressure.  

Applying  a  compression  load  at  pipe  end,  the  maximum  applied  end  moment  and  the  corresponding 
rotation decrease. For tension load up to +0.12*Ny are applied rotations and moments increase for all 
geometries, while for N=0.24*Ny no further increase is observed for PN40 case and a slight decrease 
for N=0.24*Ny.  

Tensile  strains  on  pipe  are  influenced  by  axial  loading  conditions.  This  is  particularly  true  at  pipe 
extrados where axial tensile load produce large plastic strain spots that can result in wall crack. 

II.4.3.3.2 

Parametric numerical study of non-standard flanged joints 

With  reference  to  bolted  flanged  joints,  the  aim  of  Task  4.3  was  to  use  the  numerical  FE  models  of 
calibrated bolted flange joints presented in the Deliverable D4.2, in order to understand the behaviour of 
joints for a change of flange thickness.  

A  parametric  study  of  joints  proposed  was  performed,  in  order  to  evaluate  the  different  behaviours 
under  bending  and  axial  loading.  The  flanges  used  in  Design  01  and  02  have  the  same  geometrical 
characteristics  of  the  flanges  of  Type  1,  as  indicated  in  the  standard  EN  1092-1  [23],  except  for  the 
thickness flange. Hence, a comparison among the 36 mm thickness flanges (proposed in the EN 1092-
1:2007) and the 18 and 27 mm thickness flanges (proposed by UTrento) was done. Moreover, in order 

applied rotationapplied forcesInternal pressure60 
 
 
 
 
 
 
 
 
  
  
 
 
 
 
to understand the behaviour of the joints without necks (instead of the ones suggested by the EN 1092-
1:2007,  Flanges  Type  11)  a  40  mm  thickness  flanges  was  studied.  The  results  obtained  for  different 
thickness flanges are reported in Figure II.4.3-9 for bending tests and in Figure II.4.3-10 for axial tests, 
respectively. 

600

500

400

300

200

100

0

)

N
k
(

F
d
a
o
L

Standard Flange 36 mm Type 1

Standard Flange 40 mm Type
11
Non-Standard Flange 18 mm
Type 1
Non-Standard Flange 27 mm
Type 1

0

100

200

300

Displacement Δ (mm)  

Figure II.4.3-9: Comparison among different thickness flanges for bending tests 

3500

3000

2500

2000

1500

1000

500

)

N
k
(

F
d
a
o
L

0

0

Standard Flange 36 mm Type 1

Standard Flange 40 mm Type 11

Non-Standard Flange 27 mm
Type 1
Non-Standard Flange 18 mm
Type 1

2

4

6

8

Displacement Δ (mm)  

Figure II.4.3-10: Comparison among different thickness flanges for axial tests 

In bending tests, as expected, the response of joints with thicker flanges is stronger than in non-standard 
cases. Since mainly pipes are stressed in the elastic range, FE simulations of pipes are  easier than the 
one of bolted flange joints; all numerical responses are almost identical, where only pipes are stressed. 
Furthermore,  for  the  same  load  F  applied,  the  displacement  Δ  is  greater  for  the  thinner  flanges;  this 
happens because they are subject to greater strains than the ones of standard design, and thus subject to 
major  deformations.  In  Design  01  non-standard  joint,  as  shown  in  Figure  II.4.3-11,  pipes  are  more 
stressed  than  flanges,  and  subject  to  a  buckling  phenomenon.  Instead,  in  Type  01  standard  joint,  see 
Figure II.4.3-12, the thickness increases, so flanges are more stressed; these stresses are transferring by 
bolts that result the weakest elements. In axial tests, at the same load the opening between the flanges is 
reduced in the joints with the greater thickness; this reason explains because there are different slopes of 
curves  in  the  elastic  range.  This  behaviour  is  related  to  the  fact  that  the  thicker  flanges  are  able  to 
absorb and distribute stresses in a more favourable manner, thus decreasing deformations. The standard 
joints are  therefore  able  to  withstand  more  axial  loads.  In  Design  01  non-standard  case  pipes  show  a 
buckling  phenomenon  along  the  contact  circumference.  Instead,  in Type  01 standard  case, stresses  in 
flanges become relevant, while pipes remain in the elastic range. 

The  Stress  Modified  Critical  Strain  (SMCS)  fracture criterion  was  based  to  assess  whether  the  limits 
imposed by the ABAQUS convergence were realistic with respect to the limits imposed by a reliable 
failure criterion. The critical plastic strain εp

crit can be expresses as: 

61 
 
 
 
 
 
 
 
 
 
 
 
             (     

)     

where  α  is  a  function  of  steel  properties  that  is  related  to  a  critical  void  growth  rate,  and  gives  an 
indication of the material resistance to ductile crack initiation,  σm is the mean normal stress and σeq is 
the effective flow stress. The triaxiality is defined as the ratio σm/σeq. In our case, the parameter is equal 
to 2.59 for steel pipes and flanges (S355), and equal to 5.00 for bolts (Grade 8.8). The application of the 
failure criterion was capable of identifying some elements that were deemed to be critical – as shown in 
Figure  II.4.3-13  and  Figure  II.4.3-14,  i.e.  the  elements  that  were  most  susceptible  than  others,  and 
therefore  subject  to  ductile  failure  upon  reaching  a  certain  load  limit.  The  results  obtained  from  the 
application of the SMCS failure criterion are in agreement with experimental data. In particular, Table 
II.4.3-7  and  Table  II.4.3-8  show  the  most  critical  components  connected  to  the  flange  thickness  for 
bending  and  axial  tests,  respectively.  A  careful  reader  can  notice  that  a  thickness  increase  entails  a 
failure shift from pipes to bolts. 

Table II.4.3-7: Fracture load for bending tests with SMCS criterion 

Fracture load (kN) 

Test type 

Design 

Bending 

Bending 

Bending 

Bending 

Design 01 

Design 02 

Type 01 

Type 11 

Pipe 

464.41 

476.52 

- 

- 

Flange 

485.46 

510.90 

- 

- 

Bolt 

501.11 

521.21 

476.30 

471.30 

Failure ele. 

Pipe 

Pipe 

Bolt 

Bolt 

Table II.4.3-8: Fracture load for bending tests with SMCS criterion 

Fracture load (kN) 

Test type 

Design 

Axial 

Axial 

Axial 

Axial 

Design 01 

Design 02 

Type 01 

Type 11 

Pipe 

2274.00 

2633.01 

Flange 

2367.90 

- 

- 

2886.57 

3510.68 

- 

Bolt 

Failure ele. 

2351.60 

2779.29 

1853.43 

2150.00 

Pipe 

Pipe 

Bolt 

Bolt 

(a) 

(b) 

Figure  II.4.3-11: Stress distribution  for the bending test: (a) 18 mm thickness joint,  and (b) 36 mm thickness 
joint 

62 
  
  
   
 
 
 
 
 
 
 
 
(a) 

(b) 

Figure II.4.3-12: Stress distribution for the axial test: (a) 18 mm thickness joint, and (b) 36 mm thickness joint 

(a) 

(b) 

(c) 

Figure II.4.3-13: Critical elements for the bending test: (a) pipe, (b) flange, and (c) bolt 

(a) 

(b) 

(c) 

Figure II.4.3-14: Critical elements for the axial test: (a) pipe, (b) flange, and (c) bolt 

II.4.3.4.

Parametric numerical study of Tee pipe joints 

Following  the  calibration of  the numerical  models  (Task4.2)  with  the experimental  test results  of the 
medium scale tests of critical components (Task3.5) a parametric study has been conducted. For the tee-
junction  connections  the  following  parameters  are  considered  of  importance:  (a)  tee-junction 
configurations, with a reinforcement plate or with a fitting, Figure II.4.3-15, (b) load direction, in-plane 
and  out-of-plane  bending  and  axial  loading  of  the  branch,  (c)  internal  pressure  variation,  (d)  fitting 
radius,  (e)  geometric  parameters,  diameter  and  thickness  of  branch  and  main  pipe,  (f)  reinforcement 
plate geometry (width and thickness variations). 

Finite element models have been developed in ABAQUS version 6.11 for each specimen configuration 
for the tee-junctions with a reinforcement plate and fitting. 8-noded reduced integration solid elements 

63 
 
 
 
 
 
 
 
 
 
 
 
(C3D8R) were used. Three elements through the thickness are considered adequate to capture the stress 
and strain variation through the thickness. The main pipe ends were fixed and a prescribed displacement 
was applied at the branch end. The distance from the branch to the main pipe and was at least 3 times 
the pipe diameter so that the effect from the supports was negligible. Both isotropic hardening and LKH 
material behaviors were considered for the monotonic analyses.  

Figure II.4.3-15: Reinforcement pad (EN13480-3 [25]) and tee-fitting geometry (EN10253-2 [22]). 

II.4.3.4.1 Effect of loading 

The capacity of the tee-junctions in various loading directions and internal pressure levels was studied. 
Axial loading, in-plane and out-of-plane bending of the branch was considered. The internal pressure 
values selected were 0, 30 and 60 bar. Isotropic material hardening and the full junction configuration 
were considered.  

The  capacity  in  the  axial  direction  is  larger  than  bending,  with  out-of-plane  bending  being  the  least 
strong. The internal pressure has no effect in the response of the junction up to yield but increases the 
ultimate capacity. The effect is more pronounced for axial loading rather than bending.  

A  comparison  of  the  results  of  the  two  different  tee-junction  configurations  (see  appendix  Figure 
III.4.1-18 and Figure III.4.1-19) shows that the tee junction with the fitting is considerably stronger than 
the tee-junction with the reinforcing plate in all loading direction. In addition, the deformability of the 
junction  is  lower  compared  to  the  junction  with  the  reinforcing  plate,  except  the  case  of  in-plane 
bending. There is 93% difference in the out-of-plane loading capacity, whereas the difference drops to 
55% in the in-plane bending capacity. In the axial direction the difference is more pronounced when the 
branch  is  being  pushed towards the  main  pipe. The fitting  is 22%  stronger  in  the  axial  “pulling”  and 
48% stronger in the axial “pushing” than the reinforced tee-junction. There is also big stiffness variation 
between the two configurations for the various loading directions. The stiffness of the fitting is twice 
the  stiffness  of  the  reinforced  tee-junction  in  the  out-of-plane  bending  and  is  also  significant  for  in-
plane bending. Finally, there is no significant change in stiffness between the two configurations when 
loading is axial to the branch. 

II.4.3.4.2 Effect of fitting radius 

Standard EN10253-2 [22] for pipe welded pipe fittings specifies that the crotch radius is less or equal to 
¾ of the height of the branch (rc≤0.75(G-D/2)) or 0.5 of the height (rc≤0.5(G-Ds/2)), Figure II.4.3-15. 
Because  no tolerance is provided  but  rather  guidance,  the effect  of  radius  on the  fitting  capacity  was 
studied here. Considering the limitation of ¾ of branch height the radius was varied between 35, 40 and 
45 mm. Only slight changes in the overall capacity of the tee-junctions are observed in terms of load 
resistance with the variation in radius of curvature considered. 

asaplsplpcplfscseellApAApf,,264 
 
                
 
 
 
 
 
II.4.3.4.3 Effect of branch to pipe diameter 

Three fitting geometries, with varying branch to pipe diameters were compared; an 8”x4”, 8”x6” and 
8”x8”  SCH40  configuration.  The  dimensions,  as  they  appear  in  Figure  II.4.3-15,  were  taken  from 
EN10253-2 [22]. The radius of curvature for the fitting was 30, 40 and 50 mm respectively for the three 
geometries of 8”x4”, 8”x6” and 8”x8” based on the standard maximum allowed curvature. Out-of-plane 
and in-plane bending as well as axial loads in the direction of the branch were applied. Internal pressure 
of 60 bar was applied in all cases. Linear kinematic hardening material behavior was assumed.  

The behavior of the three configurations varies greatly in all loading directions both in capacity as well 
as mode of failure. The capacity of the 8x6 fitting is 1.4 times greater than that of the 8x4 fitting, and 
the  8x8  fitting  has  a  capacity  84%  greater  than  that  of  the  8x6  fitting.  For  in-plane  bending  the 
comparative  figures  become  1.6  and  100%  respectively.  The  yield  and  ultimate  strength  variation  is 
shown in Figure III.4.1-20. 

In the case of axial loading, when the branch is pulled away from the junction fracture is expected to 
occur at the branch. The strength of the junction increases as the branch diameter increases. Whereas, 
when the branch is pushed towards the fitting the change in capacity is affected by the strength of the 
branch, the pipe and the fitting and the mode of failure changes in each configuration. For the 8”x8” 
tee-junction  failure  occurs  at  the  main  pipe,  for  the  8”x6”  tee-junction  failure  occurs  at  the  fitting, 
while, for the 8”x4” tee-junction failure occurs at the branch.  

II.4.3.4.4 Effect of reinforcing pad geometry 

For  the  8”x6”  SCH40  configuration  and  assuming  60  bar  internal  pressure  a  number  of  possible  pad 
geometries have been selected based on the designed criteria of EN13480-3, Figure II.4.3-15. The width 
and  thickness  varied  between  the  minimum  and  maximum  allowable  values.  Namely,  the  thickness 
varied between 4, 6, and 8 mm for pad width of 40mm and for a thickness of 6 mm the pad width varied 
between 20, 30 and 40 mm. 

The  effect  of  thickness  is  more  significant  than  the  width  of  the  pad  in  terms  of  bending  strength  as 
shown in Figure III.4.1-21.  When the branch is pulled axially away from the pipe there is no significant 
effect  of  thickness  and  width  of  the  pad  for  most  cases  studied.  Similarly,  when  pushing  the  branch 
towards  the  pipe  the  effect  of  pad  geometry  is  not  clear.  However,  when  the  maximum  allowed 
thickness  of  the  pad  is  used,  being  equal  to  that  of  the  pipe  thickness,  the  axial  positive  capacity 
decreases. This happens because the reinforcement is too stiff and stress concentration appears in the 
centerline of the main pipe. On the other hand, in the negative axial loading, the significantly stronger 
reinforcement pad increases strength and reduces the instability of the branch which causes buckling in 
the branch in all other four pad geometries modeled. 

II.5.  Work  package  5  –  Seismic  behaviour  and  special  seismic  design  issues  in 
liquid storage tanks; development of relevant design guidelines/recommendations 

The  main  objective  of  this  work  package  (WP5)  is  the  development  of  a  complete  set  of  design 
guidelines / recommendations, accompanied by design examples, for the seismic analysis and design of 
industrial  liquid  storage  tanks.  The  guidelines  are  compatible  with  the  existing  Eurocode  design 
framework, and extend the existing rules of Eurocode 8 (EN 1998-4).  

II.5.1. 

Task 5.1 Special issues in seismic behaviour of liquid storage tanks 

The work conducted in Task 5.1 consisted on examining several special issues on the seismic response 
of liquid storage tanks. The results of this investigation  can be considered as a background document 
for  the  seismic  design  guidelines  of  liquid  storage  tanks  developed  in  Task  5.2.  In  particular,  the 
following issues have been addressed in detail:  

(a) 

the  number  of  sloshing  modes  to  be  considered  in  convective  action  and  their  combination 
with impulsive action; 

65 
 
 
 
Three  tanks  under  consideration  have  been  subjected  to  base-ground  seismic  acceleration  from  10 
artificial  earthquakes,  generated  from  the  EN  1998-1  design  spectrum.  The impulsive  and  convective 
forces  have  been  calculated  for  each  time  increment of  the accelerogram.  The maximum  forces  have 
been compared then to each other to see how important the convective forces are to the total action and 
to  determine  how  many  sloshing  modes  are  necessary  to  be  considered  in  an  analysis.  The  two 
maximum forces were also combined with two ways: a) with the addition of the maximum convective 
and maximum impulsive forces b) the SRSS rule and compared with the maximum of the addition of 
the  two  forces.  The  analysis  aimed  at  determining  which  of  the  two  ways  of  combination  is  more 
accurate. 

(b) 

the effects of shell deformation on the seismic response;  

Steel tanks are not rigid, they deform due to their small thickness with respect to their diameter. A finite 
element analysis, which accounted for the dynamic interaction between the liquid and the deformable 
tank, has been reported by Scharf [51], where it is assumed that the motion of the tank-liquid system is 
the  sum  of  three  contributions,  referred  to  as:  ‘rigid  impulsive’,  ‘sloshing’  and  ‘flexible’.  The  third 
component  of liquid  motion  satisfies  the  condition  that  the  radial  velocity  of  the  fluid along  the  wall 
equals the deformation velocity of the tank wall, as well as the conditions of zero vertical velocity at the 
tank bottom and zero pressure at the free surface of the fluid. More simplified models have also been 
developed  for  deformable  liquid  containers,  such  as  the  Veletsos  model  [54].  Similar  to  Scharf’s 
formulation,  the  Veletsos’  model  further  assumes  that  the  convective  motion  of  the  liquid  remains 
unaffected  by  tank  shell  deformation. Therefore, the hydrodynamic  solution  for  the  rigid  container is 
still applicable, and the tank shell deformation affects only the impulsive motion. This methodology has 
been adopted by the API 650 provisions, whereas the simplified methodology for deformable containers 
in section A.3.2.2 of EN 1998-4 is in-line with this approach. The accuracy of this approach has been 
examined. 

(c) 

calculations of hoop hydrodynamic stresses;  

Thickness  design  of  liquid  storage  tanks  is  based  on  hydrostatic  pressure  and  the  corresponding 
allowable stress of the tank material, as described in API 650. During an earthquake event, additional 
hoop stresses develop due to hydrodynamic loading, which need to be added to the stresses from the 
hydrostatic  pressure  and  compared  with  the  allowable  stresses.  A  methodology  for  calculating  hoop 
stresses has been proposed in the paper of Wozniak & Mitchell [55] based on the work of Veletsos [54]. 
Consideration of this methodology for hoop hydrodynamic stresses is included in the new Appendix E 
of  API  650,  2007  edition  [3].  On  the  other  hand,  there  is  no  such  methodology  in  EN  1998-4  [17]. 
Using finite element models of three tanks, the hoop hydrostatic and hydrodynamic stresses have been 
computed and compared to the ones predicted by API 650. 

(d) 

uplifting behaviour of unanchored tanks; 

In  many  practical  applications,  relatively  broad  aboveground  liquid  storage  tanks  are  constructed 
unanchored, in the sense that their bottom plate is in simple contact with the ground without anchors. In 
such  a  case,  under  strong  seismic  loading,  the  tank  may  exhibit  uplifting  of  its  base  plate,  when  the 
magnitude of the overturning moments exceeds a threshold value. Although uplift does not necessarily 
result  in  the  collapse  of  the  tank,  its  consequences  may  lead  to  serious  damage  to  any  piping 
connections  that  are  incapable  of  accommodating  severe  vertical  displacements.  In  addition,  possible 
damage at the uplifted bottom plate may occur due to its separation from the ground and an increase in 
the axial stress acting on the tank wall part close to the ground, leading to a precipitation of “elephant’s 
foot” buckling. For this purpose, base uplifting mechanics is examined numerically through a two-step 
methodology: a) a detailed finite element shell model of the tank for incremental static analysis, capable 
of  describing  the  state  of  stress  and  deformation  at  different  levels  of  loading  and  b)  a  simplified 
modeling of the tank as a spring-mass system for dynamic analysis, enhanced by a nonlinear spring at 
its base to account for the effects of uplifting. Three cylindrical liquid storage tanks of different aspect 
ratios are modeled and examined both as anchored and unanchored. The results are aimed at comparing 
with the relevant seismic design provisions of EN 1998-4 and API 650. 

(e) 

buckling at the top of the tank; 

A sloshing motion of the tank contents occurs during earthquake motion. Observations following strong 
earthquakes indicated that the actual amplitude of liquid surface elevation may exceed several meters in 

66 
some cases. For full or near-full tanks, resistance of the roof to the impact of the sloshing wave results 
in an upward (internal) pressure distribution on the roof and buckling of the upper courses of the shell 
walls called “sloshing buckling” attributed to the alternating sign of hydrodynamic (sloshing) pressure 
on  the  tank  wall  during  the  seismic  excitation.  More  specifically,  when  –  in  the  course  of  a  seismic 
event  –  pressure  becomes  negative,  i.e.  directed  inwards,  it  may  overtake  the  internal  (outward) 
hydrostatic pressure, so that the thin-walled tank top is locally under external pressure leading to shell 
buckling.  Damage  may  also  occur  at  the  frangible  joints  between  walls  and  cone  roofs,  with 
accompanying  spillage  of  tank  contents  over  the  top  of  the  wall.  To  quantify  this  effect,  a  tank  is 
  for  7 
considered  and  the  impulsive  and  convective  pressures  are  computed  at 
artificial  earthquakes,  corresponding  to  a  maximum  ground  acceleration  of  0.25g.  Resistance  of  the 
thin-walled tank shell against this type of buckling can be provided by one or more ring stiffeners at the 
top of the tank. 

  and 

(f) 

elephant’s foot buckling; 

When subjected to earthquake loading, tanks may fail due to excessive compression at the bottom of the 
tank  shell.  This  is  a  typical  buckling  mode  of  a  thin-walled  shell  under  the  simultaneous  action  of 
meridional compression and hoop tension due to the presence of internal pressure; it is referred to as 
“elephant’s  foot  buckling”  and  is  considered  as  the  major  source  of  liquid  storage  tank  failure  under 
seismic  action.  Elephant’s foot  buckling  in  a  tank  has  been  examined  numerically,  with  emphasis  on 
post-buckling  strength,  through  the  consideration  of an  axisymmetric finite  element  model,  mainly  to 
investigate post-buckling behavior. 

(g) 

behavior factor of liquid storage tanks; 

than  the  values  of 

in  EN  1998-4.  The 
The  API  650  provisions  adopt  totally  different  values  of 
smaller values  of behavior factor for the convective motion in comparison with the impulsive motion 
are attributed to the fact that the convective response may not be capable at dissipating seismic energy. 
However, there exist striking differences on the values of 
for the impulsive component of the 
seismic force, which may not be reasonable. The ductility of the structural system constitutes the main 
parameter for the reduction of the seismic force. In particular, considering elephant’s foot buckling as 
the dominant mode of buckling, there is limited capability of the tank shell of absorbing and dissipating 
seismic energy due to its unstable post-buckling behavior. Communication of the authors with the API 
 adopted by the 
650 drafting committee indicated that the relatively high values of reduction factors 
American standards is based both on previous experience, as well as on other dissipation mechanisms 
such as radiation damping at the tank foundation (soil-tank interaction), the ductility of the anchor bolts, 
or  the  inelastic  behavior of  the  base  plate  in  the  case  of  uplifted unanchored  tanks.  Nevertheless, the 
unstable post-buckling behavior of elephant’s foot buckling is not in favour of a high value of behavior 
factor. 

 and 

(h) 

nonlinear sloshing effects;  

The calculation of hydrodynamic forces is based on linear liquid hydrodynamic theory, assuming small-
amplitude of sloshing waves. The question is whether this assumption is valid or not in the course of a 
strong earthquake. This requires a more refined analysis. A first analysis has been conducted by Chen et 
al. [14], for rectangular liquid storage tanks with finite elements. In addition, an asymptotic solution of 
the nonlinear problem has been reported in the recent study of Katsikogiannis [41]. In that study, a wide 
range of tank geometries have been subjected to real earthquake events. 

(i) 

floating roof behaviour. 

For the case of a Floating Roof (FR) storage tank, the seismic loads can induce a large displacement of 
the  FR  due  to  the  liquid  sloshing.  This  FR  motion  possibly  leads  to  severe  structural  damage  and 
catastrophic fire accidents as experienced in Japan and Turkey during past earthquakes. The main issues 
that  should  be  taken  into  account  in  designing  floating  roofs  in  seismic-prone  areas  are:  adequate 
freeboard,  the  stress  conditions  in  outer  pontoon,  seal  friction,  damping  due  to  seal  friction, 
modification of wall pressure due to floating roof motion, flexible nature floating roofs and effects of 
base isolation on FR motion. 

Considering  typical  liquid  storage  tanks,  provided  by  our  industrial  partners,  the  investigation  of  the 
above issues has been aimed at evaluating EN 1998-4 provisions (in comparison with those of API 650) 

0RqqRR67 
for the seismic design of liquid storage tanks, towards proposing possible improvements/amendments of 
EN 1998-4.  

The main conclusions from the investigation of these tanks can be stated as follows: 

  For the majority of tanks, consideration of only one sloshing (convective) mode is adequate for 
the calculation of the total seismic force. Furthermore, the combination of the convective force 
with the impulsive force should be conducted according to the SRSS rule. 

  Using a systematic two-step methodology for analyzing the seismic response  of uplifting tanks 
has shown that the EN 1998-4 provisions provide a reasonable yet conservative approach of the 
increase of axial compression due to uplifting. 

  The Veletsos methodology for deformable containers can be used for design purposes. 
  A simple methodology is presented in order to design tanks against buckling at the top, due to 

liquid sloshing, using the concept of “equivalent” external pressure. 

  The behavior (reduction) factor for the seismic design of liquid storage tanks is a controversial 
issue that has not been examined thoroughly so far. Because of the relatively low post-buckling 
strength of the shell tank, the authors would be quite cautious in using a behavior factor greater 
than 1.5 for the impulsive component of motion. 

  There  exists  little  information  on  current  design  practice  for  floating  roofs;  to  avoid  complete 

tank destruction, one should avoid seal damage to prevent fire ignition. 

  Nonlinear sloshing effects on the seismic behavior have not been thoroughly investigated so far. 
Limited  numerical  investigations,  reported  elsewhere,  indicated  that  their  effects  on  the  total 
seismic  force  may  not  be  important,  but  they  can  have  significant  effects  on  freeboard 
considerations. 

II.5.2. 

Task 5.2 Design recommendations / guidelines for the seismic analysis & design 
of liquid storage tanks 

The design guidelines for the seismic design of liquid storage tanks have been developed in the course 
of Task 5.2, in an attempt to present a document that improves the existing provisions of EN 1998-4. 
The  structure  of  those  guidelines  is:  1  Introduction;  2  Relevant  Standards  and  Guidelines;  3  List  of 
Symbols; 4 Seismic Actions in Liquid Storage Tanks; 4.1 Spectrum for seismic analysis and design; 4.2 
Behavior  factor;  4.3  Importance  factor;  4.4  Calculation  of  Seismic  Forces  and  Moments;  4.4.1 
Impulsive/ convective decomposition of motion; 4.4.2 Vibration periods and damping coefficients; 4.4.3 
Seismic Hydrodynamic pressures; 4.4.4 Horizontal seismic forces; 4.4.5 Ringwall and slab overturning 
moment; 4.4.6 Combination of impulsive / convective action; 4.4.7 Vertical Seismic Action; 4.4.8 Hoop 
Hydrodynamic  Stresses;  4.4.9  Meridional  Stresses;  4.4.10  Forces  on  anchors;  4.4.11  Elevation  of 
liquid free surface; 4.4.12 Seismic Action on nozzles (pipe attachments); 5 Seismic Resistance of Liquid 
Storage Tanks; 5.1 Limit States (General); 5.1.1 Shell Buckling Mode; 5.1.2 Anchorage Failure; 5.1.3 
Roof Damage; 5.1.4. Hoop Hydrodynamic Stresses; 5.1.5 Base Plate Connection Failure; 5.1.6 Failure 
at Attached Piping Locations; 5.2 Tensile Strength of Tank Shell; 5.3 Tank shell buckling; 5.4 Stability 
against sliding; 5.5 Design of Nozzles; 5.6 Resistance of Anchorage; 5.7 Sloshing; 5.8 Foundation; 6 
Special  seismic  design  provisions;  6.1  Bunding;  6.2  Floating  Roof  Requirements;  6.3  Constructional 
Details;  ANNEX  A:  Hydrodynamic  loading  due  to  seismic  action;  A.1  Introduction;  A.2  Horizontal 
seismic  action;  A.2.1  General;  A.2.2  Impulsive  pressure  /  resultants  /  heights;  A.2.3  Convective 
pressure  /  resultants  /  heights;  A.3  Vertical  seismic  action  in  deformable  tanks;  A.4  Sloshing  wave 
height; A.5 Soil-Tank interaction; ANNEX B: Unanchored tanks; B.1 General  – uplifting; B.2 Effects 
on meridional compression; B.3 Size of uplift and length; B.4 Base plate and deformation; ANNEX C: 
Shell  Buckling  Resistance;  C.1  EN  1998-4  Shell  buckling  provisions;  C.1.1  Elastic  Buckling;  C.1.2 
Elastic-plastic collapse; C.2 API 650 Shell buckling provisions.  

The  guidelines  can  be  employed  by  designers  of  liquid  storage  tanks  in  seismic  areas,  providing 
background information to existing standards, and identification of ultimate limit states under seismic 
action, as well as amending and improving the existing provisions of EN 1998-4. Design examples of 
three liquid storage tanks provided by our partners have been considered and designed against seismic 
loads to illustrate the use of those guidelines. These tanks are: 

68 
 
(a)  Tank I is a moderately-broad tank. This is a 27.4-meter-diameter tank with total height 16.5 
1.131). The tank thickness varies from 6.4 
meters. The filling height 
mm at its top course to 17.7 mm at its bottom course; it is unanchored with bottom plate of 6 
mm thick and an 8-mm-thick annular plate. 

is 15.7 m (

(b)  Tank II is a tall 18-meter-diameter tank with a total height of 20 meters and with filling height 
2.1). Tank thickness varies from 6 mm at its top course to 10 

 equal to 19 m (

mm at its bottom course and it is anchored. 

(c)  Tank III, is a very broad 68-meter-diameter tank with a total height of 20 meters. The filling 
0.558). Tank thickness varies from 34 mm at its top 

is equal to 19 m (
height 
course to 38.5 mm at its bottom course. 

(d)  All  three  of  the  above  liquid  storage  tanks  are  filled  with  water  (

1000  kg/m3)  and  the 
material of the tank shell, the bottom plate and the roof is structural steel S235 (equivalent to 
A36 steel) with yield stress 

235 MPa. The tanks are depicted in Figure II.5.2-1. 

Figure II.5.2-1: The three liquid storage tanks used in the design examples. 

II.6.  Work package 6 –   Seismic  behaviour  and  design  of  industrial  pressure 

vessels; development of seismic design guidelines/recommendations 

The  main  objective  of  this  workpackage  (WP6)  is  the  development  of  design  guidelines  / 
recommendations for the seismic analysis and design of industrial pressure vessels (vertical cylinders, 
horizontal  cylinders  and  spheres).  The  guidelines  will  concern  various  geometries  and  support 
conditions and will be compatible with the existing Eurocode 8 (EN 1998) framework for earthquake-
resistant design.  

II.6.1. 

Task 6.1 Seismic behaviour of vertical-cylindrical pressure vessels 

The  seismic  behavior  of  different  geometries  of  vertical  vessels  on  skirt  supports  has  been  analyzed 
using pushover and incremental dynamic analysis techniques. The failure mechanisms were discussed 
and possible failure criteria were proposed. Two regions of the vertical vessel have been identified as 
critical,  namely  the  skirt  and  the  1st  shell  course.  The  behavior  factor  q  determined  for  the  vertical 
vessels was found to be 1.5 lower than the value of 2 suggested by ASCE7 [5] for such pressure vessel 
geometries. The effect of internal pressure was insignificant up to the point of maximum capacity of the 
vessel but beneficial for post-buckling behavior. 

II.6.2. 

Task 6.2 Seismic behaviour of horizontal-cylindrical pressure vessels 

The seismic behavior of different geometries of horizontal pressure vessels on saddle supports has been 
analyzed  using  pushover  analysis  techniques.  The  effect  of  reinforcing  rings,  length,  thickness  and 
internal  pressure  was  studied  using  pushover  analysis.  The  mode  of  failure  is  buckling  at  the  saddle 
support. The exact location in the saddle varies with load direction. The effect of internal pressure was 
more significant when loading was applied in the transverse rather than in the longitudinal direction due 
to the change of failure location to the shell area adjacent to the saddle reinforcement. Reinforcing rings 
are recommended for large pressure vessels or pressure vessels located in high seismicity areas. Caution 

HHRHHRHHRy69 
 
 
 
 
 
 
shall be taken when deciding the area of the saddle reinforcement as the adjacent shell area is expected 
to fail in the case of no internal pressure. 

II.6.3. 

Task 6.3 Seismic behaviour of pressure vessels of spherical shape 

Within  this  Task  one  example  of  an  existing  pressure  vessel  supported  by  non-braced  columns  with 
circular hollow sections has been investigated, extending the work carried out in Deliverable 1.3b for 
case  study  4.  Further  investigations  were  carried  out  using  the  same  vessel  supported  by  braced 
columns as well as parametric geometries, taking into account both vessels as references. For all vessels 
modal  and  pushover  analyses  have  been  performed  with  the  aim  to  assess  the  influence  of  different 
geometrical parameters on the non-linear behaviour as well as to identify critical failure modes (limit 
states). In addition a method for calculating the fundamental period using a simplified beam model were 
derived and verified on the basis of the numerical investigations. An excellent correlation between the 
above  mentioned  methods  has  been  observed.  The  deviations  lie  below  2%  on  average  and  6.2%  on 
maximum  for  non-braced  systems  and  also  below  2%  on  average  and  8.0%  on  maximum  for  braced 
systems, whereby the maximum deviations refer to systems where the recommended limit values for the 
dimensions  are  not  complied.  In  order  to  determine  exemplary  the  behaviour  factor  q  non-linear 
dynamic  analyses  were  performed,  with  the  result  that  for  both  examples  the  behaviour  factor  was 
determined and showed to be bigger than 2.0. A behaviour factor of 2.7 was found for the non-braced 
system,  which  should  be  restricted  to  this  value  since  dynamic  instability  phenomena  can  occurs 
suddenly  using  higher  values.  For  the  braced  system  a  behaviour  factor  of  2.3  was  identified,  when 
limitation of the plastification to the bracings is desired. Otherwise higher values than 3.0 were defined, 
if  additional  plastification  of  column  head  is  allowed.  Moreover  the  parametric  numerical  study 
(pushover  analyses)  has  been  shown  that  by  compliance  of  the  limit  values  for  the  dimensions  of 
spherical  pressure  vessels  the  displacement  ductility  μ  =  eult/ey  (where  eult  and  ey  represent  the 
displacements by reaching the ultimate load and yield load respectively) for different braced and non-
braced systems with 4, 8 or 12 columns is relatively stable. For the non-braced systems slightly higher 
values than μ = 2.0 were defined, whereas for the braced systems the displacement ductility showed to 
be bigger than μ = 4.0.  

However  these  recommendations  apply  only  for  spherical  pressure  supported  by  an  even  number  of 
braced or non-braced columns with circular hollow sections as well as only on the condition that the 
limit  values  for  the  dimensions  of  the  vessels  given  within  the  guidelines  (Deliverable  D6.2)  are 
complied. 

II.6.4. 

Task  6.4  Development  of  design  recommendations  /  guidelines  for  the  seismic 
design & analysis of industrial pressure vessels 

The  developed  guidelines  contain  design  recommendations  for  industrial  pressure  vessels  in  seismic 
regions.  The  considered  types  of  pressure  vessels  are  vertical  pressure  vessels  on  skirt  supports, 
horizontal  pressure  vessels  on  saddle  supports  and  spherical  pressure  vessels  supported  by  an  even 
number of braced or non-braced columns with circular hollow sections. For each type of vessel rules for 
seismic analysis and design are provided. Furthermore recommendations for achieving ductile response 
are given by means of global design concepts and local detailing rules. The recommendations comprise 
common  rules  valid  for  any  type  of  pressure  vessels  as  well  as  rules  depending  on  the  type  of  the 
pressure vessel. They are meant to be considered in addition to existing seismic design rules for steel 
structures  which  -  to  large  extend  -  remain  valid  for  pressure  vessels  and  in  particular  for  their 
supporting structures. 

In  particular  the  determination  of  fundamental  periods,  application  of  adequate  behaviour  factors, 
determination of seismic forces taking into account effects of liquid sloshing and recommendations for 
the seismic design as well as the definition of limit states (failure modes) and dimension limits for the 
pressure  vessels  and  the  supporting  structures  are  in  the  focus  of  these  guidelines.  The 
recommendations  and  seismic  design  rules  are  also  illustrated  by  means  of  design  examples  given 
within  the  background  document  (Deliverable  D6.1).  Furthermore  some  recommendations  regarding 
the  influence  of  large  supporting  buildings  in  terms  of  modified  floor  response  spectra  and  floor 
acceleration characteristics are given. 

70 
 
The  documentation  of  the  scientific  and  engineering  background  for  these  guidelines  as  well  as  the 
extensive  representation  of  the  design  examples  are  provided  within  the  background  document 
(Deliverable D6.1), which composed of the extended work of Task 6.1, 6.2 and 6.3. 

II.7.  Work  package  7  –  Seismic  behaviour  and design  of  industrial  piping systems 
and components; development of seismic design guidelines/recommendations 

The main objective of WP7 is the development of design guidelines / recommendations for the seismic 
analysis and design of industrial piping, including their support systems. Special emphasis will be given 
to the design of an adequately dissipative support system, to reduce the seismic action on the piping. 
The guidelines will be compatible with the existing Eurocode 8 (EN 1998) framework for earthquake-
resistant design. 

II.7.1. 

Task 7.1 Seismic behaviour of industrial piping and their support systems 

II.7.1.1.

Analysis of a Representative Piping System 

Piping  systems  represent  key  structural  systems  for  an  industrial  facility.  Few  contributions  in  the 
literature  are  available,  in  order  to  clarify  the  seismic  requirements  of  piping  systems.  In  addition, 
aspects like seismic action and structural modeling have not yet been treated in a satisfactorily manner. 
Moreover, current American and European standards do not contain adequate information and relevant 
provisions for a proper seismic analysis and design of piping systems. Along these lines, some problems 
relevant to seismic analysis criteria of piping systems are addressed in this investigation. In particular, 
several aspects are treated: 1) modeling of pipes and pipe-racks; 2) selection of the analysis method; 3) 
definition  of  the  seismic  action;  4)  dynamic  analysis  of  the  system;  5)  stress  analysis  of  pipes;  6) 
definition of the ultimate capacity of pipes and joints between pipes.  

Figure II.7.1-1: The piping system considered in the present study. 

To this end, in the first part Work package 7 (WP7, Task 7.1) a representative case study of an actual 
piping  system  is  analysed,  as  shown  in  Figure  II.7.1-1.  The requirements  of  American  and  European 
codes  are  compared  and  important  aspects  are  highlighted:  a)  definition  of  numerical  models  for 
seismic  analysis,  b)  seismic  input  and  analysis  methods,  c)  piping  stress  analysis  and  checks  d) 
evaluation of the ultimate capacity of pipes and joints necessary for a correct design of a structure, as 
suggested by modern approaches like Performance-based Engineering. A discussion of the main results 
with  respect  to  experimental  data  for  a  flange  joint,  carried  out  at  University  of  Trento,  allowed 
highlighting the conservatism of current standards. 

In particular, the present investigation has been aimed at clarifying all the design steps of this type of 
structural  systems,  analyzing  the  representative  case  study  according  to  both  European  [25]  and 
American  standards  (ASME  B31.3).  The  comparison  between  EN13480  and  ASME  B31.3  standards 

71 
 
 
 
 
 
showed  that  the  two  codes  provide  similar  indications  for  the  evaluation  of  the  seismic  response  of 
pipelines  supported  by  a  pipe-rack.  The  usual  method  of  designing  piping  systems  is  based  on  the 
allowable stress approach, considering that the structure is elastic. Consequently only an operating basis 
earthquake condition (OBE) can be taken into account for design. Conversely, the modern approach to 
the  seismic  design  of  structures  is  to  differentiate  serviceability  from  ultimate limit  states. This  latter 
condition  could  be  represented,  for  example,  by  the  Safe  Shutdown  Earthquake  (SEE).  However, 
appropriate  definition  of  the  deformation  capability  of  the  pipes  and  fittings  (ovalization,  etc.)  in  the 
inelastic range would be necessary. The advantages of such an approach would be several and would 
lead to design optimization. For example in the case study under consideration, seismic action results in 
a bending moment far below the moment capacity of a bolted flange joint located in the piping system. 
This  conservatism  seems  to  be  in  contrast  with  the  modern  performance  based-design  approach,  for 
which a certain level of yielding in the structure may be allowed according to a specific performance. 
For instance, it would be possible to accept a certain level of yielding of pipes under the condition that 
leakage would not occur, leading to a reduction of the thickness of the flange and, therefore, at a more 
efficient and economical design. 

II.7.2. 

Task  7.2  Development  of  design  recommendations  /  guidelines  for  the  seismic 
analysis & design of industrial piping systems 

The  second  part  of  work  package  7  (WP7, Task  7.2)  is  dedicated to  the  development  of  guidelines 
for  the  seismic  design  of  industrial  piping  systems that synthesize the work conducted in the previous 
work packages on the same topic (WP1, WP2, WP3 and WP4). The  guidelines  consist  of  a  concise 
document compatible  with  the  existing  Eurocodes  design  framework,   which  includes some  useful 
indications on seismic analysis of piping systems. The developed document follows the basic principles 
of  earthquake  engineering  expressed  in  EN1998:1, extending  the  existing  rules  to  industrial  piping 
components  and  systems.  The  guidelines  are  meant  to  be  used  by  designers  of  industrial  piping 
systems and provide: 

  background  information  to  the  existing  standards,  including  calculation  examples,  identifying 

differences between these standards and evaluation of these differences; 

  identification of critical steps and parameters in seismic design of piping systems;  
  recommendations for modeling of piping systems for seismic analyses; 
  identification  and  calculation  example  of  the  interaction  between  piping  systems  and  the 

supporting substructure; 

  identification  of  bearing  resistance  (ultimate  limit  states)  for  piping  components,  as  obtained 
from experimental programs , performed within the framework of the INDUSE project,  such as 
bends (elbows), tee-joints , flanges and tank nozzles, including recommendations for the use of 
these components in piping systems, subject to seismic action.  

The  guidelines  are  organized  as  follows:  Scope;  Definitions;  Reference  documents,  Symbols  and 
Abbreviations,  Methods  and  relevant  parameters  for  seismic  analyses;  Seismic  input;  Modeling  of 
piping  systems;  Dynamic  seismic  analyses;  Limit  states  (including  results  obtained  from  the  testing 
programs);  Recommendations.  The  use  of  these  guidelines  is  explained  in  appendices,  presenting 
calculation examples on: 

A. seismic analyses of different types of piping systems 
B. interaction between piping systems and the supporting pipe rack 
C. calculation  of connections between piping and tank wall and selection of nozzle types   
D. Example  for load combinations and load factors for ultimate lmit states, using the LRFD method 

Flow charts are also developed (see Appendix II.7), to facilitate their use. 

The Guidelines deals with the results of  a comparative study of the relevant European and American 
Codes and Guidelines investigating possible  differences in seismic input (elastic response spectra) are 
evaluated  and  conclusions  drawn  (Appendix  II.7).  A  direct  comparison  between  the  equivalent  static 
analyses  versus  linear  dynamic  analyses  indicated  that  the  uniform  load  seismic  calculations  yielded 
more conservative results than the dynamic seismic calculations, provided the natural frequencies are 
high enough. However, for long and slender piping systems, a dynamic analysis may be of fundamental 
importance  especially  if  the  natural  frequency  is  lower  than  4-5Hz.  In  such  cases,  the  uniform-load 

72 
method  might  substantially  underestimate  the  stresses  and  forces.  The  design  of  the  piping  system  is 
more important than the differences between codes in seismic design spectra. This is the decisive factor 
on  whether  the  system  is  capable  of  withstanding  a  particular  seismic  event,  resulting  in  a  trade-off 
between satisfying the requirements of static and dynamic analysis. 

II.7.2.1.

Influence of Supporting Structure 

The  results  of  earlier  work  packages,  dealing  with  dynamic  coupling  between  piping  system  and  the 
supporting  structure  (pipe  rack)  are  evaluated  in  the  design  guideline,  giving  suggestions  for  the 
behavior factor for a moment-resistant pipe-rack frame, as well as a criterion based on the WR ratio on 
whether the pipe weight and stiffness should be taken into account. This is illustrated in the calculation 
example in Annex B of the document (Appendix II.7). 

II.7.2.2.

Recommendations for Piping Modeling  

When performing dynamic analyses it is very important to apply the correct modeling practice in order 
to  obtain  reliable  results,  however  this  is  not  addressed  in  the  present  codes.  The  Guidelines  provide 
recommendations for modeling, in order to obtain the required calculation accuracy, as well as point of 
attention  when  using  commercially  available  piping  software  for  dynamic  analyses.  Also  some 
guidance is given for the design forces on restraints and the influence of gaps. It is advised to include 
those recommendations for modeling in the next revision of EN 1998-3 and ASCE 7-05 

II.7.2.3.

Limit States of Industrial Piping  

Available  piping  codes  and  standards,  such  as  ASME  B31.3  and  EN  13480-3,  appear  to  be  rather 
conservative for the case of seismic design. The results of literature study have been combined with the 
results of extensive numerical analyses of piping components and validated against extensive testing of 
piping components and parametric analyses. From this evaluation, proposals for limit states of piping 
components, such as bends (elbows), Tee-junctions, flanges, and tank nozzles have been developed and 
further recommendations for the use of these components are given. 

II.8.  Work package 8 – Coordination of the project 

The  main  objectives  of  the  WP8,  as  described  in  Tasks  8.1,  8.2,  8.3  and  8.4  were  to  establish  an 
efficient management process, to define standard protocols, to carry out both a SWOT evaluation and 
the monitoring of activities of the project, towards successful completion. 

II.8.1. 

Task 8.1 Efficient management process 

INDUSE  project  comprises  several  partners  and  specific  objectives.  The  co-ordinator  (University  of 
Thessaly)  has  been  responsible  to  manage  the  project  activities  to  complete  successfully  on  time.  In 
order to manage the project work certain necessary steps were taken. 

(i) 

The  representatives  of  each  partner  meet  once  in  approximately  every  six  month  for  work-
progress,  discussing  the  technical  issues  of  the  project,  and  planning  the  next  activities.  The 
meetings provided an opportunity for formal discussion in detail among partners. Seven meetings 
were held among partners of the project during the period of concern, i.e. 01 July 2009-31 Dec 
2012. 

(ii)  An  FTP server  directory  (ftp://ftp.mie.uth.gr/)  was  created  by  the  University  of  Thessaly.  Each 
partner could access this directory and it could download and upload necessary information and 
documents related to the project. The documents related to the project work, i.e. research papers, 
thesis report etc., the agenda and minutes of the meeting, experimental as well as simulation data 
and reports were uploaded on the server. This provided an efficient distribution of the urgent and 
necessary information helpful to the progress of the project. 

(iii)  A  website  of  the  INDUSE  project  (www.mie.uth.gr/induse)  was  created  by  the  University  of 

Thessaly, to disseminate the results of the project. 

73 
 
 
 
 
 
 
(iv)  The  co-ordinator  and  all  the  partners  were  in  regular  contact  via  e-mail.  The  necessary  issues 
were also discussed by phone or Skype. E-mail was the communications medium that was used at 
the  large  and  preliminary  scale.  If  more  than  two  partners  were  involved  in  discussion, 
teleconferencing was used. 

Partners  who  were  conducting  similar  activities,  for  example  RWTH  Aachen  and  EBETAM  in  the 
course of WP6, discussed together technical and administration issues several times by teleconference. 
This helped for the smooth progress of the work. For very critical and urgent issues two partners for 
example  UThessaly  and  EBETAM,  met  together  in  Volos  in  order  to  discuss  the  production  of 
specimens. UThessaly and Technip have also met several times in Athens. 

II.8.2. 

Task 8.2 Definition of the standard protocols 

A protocol is a document that describes the objective(s), design, methodology, statistical considerations 
and organization of a trial. In the course of INDUSE project, the coordinator, together with partners, set 
standard protocols for the tests regarding specimen order/manufacturing and test performance. INDUSE 
project involved several types and number of laboratory tests to be performed by the different partners. 
It is necessary, in this situation to follow a common set of procedures to prepare the tests. This reduces 
the  discrepancy  in  the  test  results  as  well  as  increases  the  efficiency.  The  following  steps  have  been 
taken in this respect: 

In particular, 

  The pipe material for all specimens (straight pipes and fittings) was decided to be P355, which is 

equivalent to X52. 

  Elbow and flange specimens have been decided to be 8-inch pipes.  
  Nozzles have been decided to correspond to 6-inch pipes 
  Tee-junction specimens were 8-inch-to-6-inch pipe branches. 
  Elbows  (to  be  used  in  the  experiments  of  3  different  partners)  were  provided  by  the  same 

manufacturer. 

  The flange and Tee used in the pseudo-dynamic loading test, has been the same as the ones used 

in the individual tests in CSM and UThessaly respectively. 

  The  final  drawings  of  the  specimens  by  the  manufacturer  were  provided  and  verified  by  the 

different partners to ensure compatibility between the users and the manufacturer. 

  Several analyses have been performed in order to examine the ultimate states of the specimens 

and compare with the testing capabilities of each partner. 

  The  welding  procedure  specifications  (WPS)  were  chosen  for  the  entire  set  of  specimens 

following the same criteria. 

  The test protocols regarding performance of tests were made and the information in detail about 
the types of tests, loadings, number of cycles, strain rate and instrumentation were described in 
WP3.  

II.8.3. 

Task 8.3 Monitoring of project activities and evaluation of the project 

The monitoring of activities was regularly performed through the co-ordination meetings in about every 
six month period. Seven meetings were held among partners of the project during the period of concern 
(01 July 2009 – 31 Dec 2012). 

In the kick-off meeting held in Athens, Greece, on 24-25 September 2009, the objectives of the project 
was  recalled  together  with  the  distribution  and  content  of  the  work  in  different  WPs  related  to  the 
concerned  period.  The  case  studies  in  WP1  were  also  chosen,  some  preliminary  discussion  on  the 
experiments was made, and planning work in WP2 was presented. In the second meeting held in Rome, 
Italy,  on  21-22  January  2010,  the  design  and  analysis  of  the  case  studies  were  discussed  and  issues 
related to preparation of specimens were discussed. In the third meeting held in Delft, The Netherlands, 
on 08-09 July 2010, the specimens geometry were finalized and the set-up has been discussed in terms 
of several initial numerical simulations. In the fourth project meeting, held in Trento, Italy, on 20-21 
February 2011, the first experimental results were conducted, and the results from WP1 and WP2 were 
summarized, for the development of the corresponding deliverables D1 and D2. The fifth meeting was a 

74 
 
short one, held in Corfu, Greece, on 25 May, 2011, in the course of the COMPDYN 2011 conference. 
In  that  meeting,  a  brief  overview  of the  progress  (mainly  the experimental  work)  has  been  made  and 
several  papers  were  presented  by  the  partners  in  dedicated  conference  sessions.  In  the  sixth  project 
meeting,  held  in  Aachen,  Germany,  on  23-24  September  2011,  the  progress  of  the  experimental  and 
numerical results has been discussed and a first discussion on the design guidelines has been made. In 
the  seventh  project  meeting,  held  in  Volos,  Greece,  on  23-24  April  2012,  apart  from  summarizing 
experimental  and  numerical  work,  the  design  guidelines  in  WP5,  WP6  and  WP7  were  extensively 
discussed.  Furthermore,  the  improved  simulation  models  of  piping  components  were  presented  and 
discussed. Finally, in the last (eight) project meeting, held in Thessaloniki, Greece, on 6-7 November 
2012, the entire experimental results and the corresponding numerical simulations were discussed and 
the  guidelines  under  preparation  has  been  extensively  presented  by  the  corresponding  partners; 
furthermore,  the  format  of  the  draft  final  report,  the  deliverables  and  the  contents  of  the  design 
guidelines were presented and discussed. 

A  SWOT  analysis  has  also  been  conducted  for  evaluating  the  Strengths,  Weaknesses,  Opportunities, 
and  Threats  involved  in  the  INDUSE  project.  It  involves  specifying  the  objective  of  the  project  and 
identifying  the  internal  and  external  factors  that  are  favorable  and  unfavorable  for  achieving  the 
prescribed objectives. The SWOT analysis of the INDUSE project can be outlined as follows: 

Strengths 

(i)  A good partnership between universities, research centres and industrial partners,  
(ii)  A multi-discipline expertise from structural, mechanical and material engineering, 
(iii)  Earlier experience of most of the partners on similar (European) projects,  
(iv)  Very good facilities for experimental testing,  
(v)  Excellent connections of EBETAM, CSM and TU Delft with pipe manufactures  
(vi)  Selection of case-studies facilitated by the participation of Technip and Tebodin. 
(vii)  Experience  on  manufacturing  of  steel  piping  components  by  CSM,  EBETAM,  TU  Delft  and 

RWTH Aachen. 

Weaknesses 

(i) 

The  coordinator  had  to  manage  a  large  consortium  (7  partners  and  1  subcontractor),  with 
different specializations, expertise and background. 

(ii)  European  specifications  for  industrial  facilities  (tanks,  pressure  vessels  and  piping)  have 

several limitations, and have not been used in everyday design practice 

(iii)  Harmonization  of  experimental  activities,  in  terms  of  specimen  size,  geometry,  loading 

schemes.  

(iv)  Ordering  of  material;  fittings  and  pipes  from  P355  material  was  not  easy  to  find  in  pipe 

manufacturers, especially for the small quantities required in the present project. 

Opportunities 

(i) 

(ii) 

INDUSE is a pioneering project of its kind, offering excellent opportunity for state-if-the-art 
scientific work.  
INDUSE is a great opportunity to develop novel and unique design guidelines, very helpful 
for current design practice. 

Threats 

(i)  A small delay of several experiments has delayed the numerical work and, by consequence, it 

delayed the preparation of the guidelines. 

(ii)  The difficult adaptation of several rules from API and ASME standards to the EN framework 

of the present guidelines. 

II.8.4. 

Task 8.4 Progress, Mid-Term and Final reports 

The following reports have been prepared and uploaded to CIRCABC: 

 
 
 
 

1st annual report – March 2010. 
2nd Mid-Term report – March 2011. 
3rd annual report – March 2012. 
4th Final report – March 2013. 

75 
The deliverables of the project have also been uploaded to CIRCABC (see list of deliverables below). 
In addition, minutes of each meeting have been held, and are available upon request. 

II.9.  Conclusions 

INDUSE

  project  has  been  a  pioneering  project  within  the  RFCS  research  program.  Combining 
The 
experimental,  numerical  and  analytical  tools,  high-level  research  has  been  performed  in  the  inter-
disciplinary  area  of  seismic  safety  of  steel  industrial  facilities.  Extensive  experimental  and  numerical 
work  has  been  conducted,  to  investigate  the  mechanical  behavior  of  critical  steel  components;  their 
behaviour under strong cyclic loading has been shown well above the requirements of existing design 
the  development  of  design 
provisions. 
guidelines/recommendations (a) expanding EN 1998-4 provisions towards an integrated seismic design 
of liquid storage tanks and (b) extending the applicability of Eurocode 8 seismic design concepts for the 
cases of industrial pressure vessels and attached equipment and piping systems including their support 
systems.  

In  addition, 

  project 

INDUSE

resulted 

the 

in 

The  project  has  been  executed  smoothly,  without  any  major  delays  or  adjustments,  and  successfully. 
The results and deliverables of 

 are both novel and unique, resulting in  

INDUSE

 

 
 

better understanding of structural behaviour of steel tanks, pressure vessels and piping under 
strong cyclic (seismic) loading,  
more reliable definition of ultimate limit states for critical industrial components,  
significant improvement of seismic design state-of-the-art of industrial facilities, through the 
developed design guidelines, 

and, most importantly,  

 

offering  the  necessary  scientific  documents for  Code  Drafting  Committees for  improving 
existing design standards within the EN design framework, towards safer design of industrial 
power plants and chemical/petrochemical facilities. 

II.10.  Exploitation and impact of the research results 

INDUSE

 has been the development of design guidelines for the seismic design of 
The main outcome of 
industrial tanks, PVs and piping systems. Those guidelines have been developed in close cooperation 
with  our  end-user  partners,  Technip  and  Tebodin  (subcontractor  of  TU  Delft),  improve  the  design 
provisions  for  industrial  tanks,  PVs  and  piping  systems  that  exist  within  the  EN  framework,  and  are 
expected to contribute significantly in more reliable seismic design of such structural systems.  

Furthermore,  the  scientific  research  in  this  topic,  associated  with  state-of-the-art  experimental  and 
numerical work in the research laboratories of the partners, has improved significantly the know-how at 
  is  reflected  in  the  numerous 
an  international  level.  The  high-quality  of  the  research  within 
scientific  publications  of  the  research  group,  and  the  special  sessions  organized  in  COMPDYN  2011 
and PVP 2013 conferences.  

INDUSE

As  a  result,  it  is  believed  that  all  the  research  work  conducted  in 
innovative, and has made a significant impact in both the scientific and the professional community. 

  has  been  novel  and 

INDUSE

76 
 
III.  Appendices 

III.1.  Appendix of WP1 

III.1.1.  Task 1.1: Collection and evaluation of available relevant experimental data 

III.1.1.1.

Experiments on piping components and piping systems 

Tests indicated that the main failure mechanisms  are (a) pipe collapse, (b) pipe low-cycle fatigue and 
(c) pipe hoop ratcheting. All tests indicated that there is some significant attenuation of due to inelastic 
material  response,  resulting  in  inelastic  energy  dissipation.  It  should  be  noted  that  the  piping  design 
specifications  are  based  on  elastically  predicted  stresses.  Therefore,  the  elastically  predicted  stresses, 
strains, accelerations, and support loads from the piping analysis for the SSE seismic level will be over 
predictions of the actual response quantities. The attenuation of elastically predicted response quantities 
are demonstrated by these tests.  

It  is  noted  that,  for  the  interpretation  of  tests  results,  where  material  properties  of  tested  components 
were  measured  and  reported,  the  actual  measured  yield  stress  has  been  significantly  above  the  code 
minimum value in most cases. Therefore, it is reasonable and appropriate to evaluate the load level in 
terms of the material yield stress not the code minimum value. In these test programs different methods 
are  used  to  specify  the  intensity  of  the  dynamic  input.  It  appears  more  appropriate  to  specify  the 
intensity of the seismic input in terms of the elastic response.  

The  tests  results  indicate  a  substantial  conservatism  of  design  provisions  for  dynamic  response  of 
piping. As shown by these tests, the dynamic primary bending stress can significantly exceed the 2Sy 
limit  adjusted  for  actual  material  properties  without  a  collapse  failure.  However,  the  amount  of 
conservatism  in  the  limit  for  all possible  piping  configurations  may  not  be  possible to  establish from 
those tests and propose a relevant amendment of the specification limit.  

It also noted that almost all the tests have been conducted on piping with a frequency of 4 Hz or greater. 
The  tests have  demonstrated  that  piping  frequency  has  a  significant effect on response;  the lower the 
frequency, the greater the response. Therefore, direct applicability of the test results is limited to piping 
with a frequency of 4 Hz or greater. The test data clearly shows the superiority of stainless steel piping 
in comparison to carbon steel piping in terms of the cyclic life (fatigue failure mode) for dynamic loads. 
However,  for  large  deformation  or  collapse  modes  of  failure,  stainless  steel  may  not  be  superior  to 
carbon steel.  More discussion of the failure mode is offered in the paragraphs below. 

Finally,  the  safety  margins  (factors  of  safety)  indicated  that  the  above  component  and  piping  system 
tests  need  to  be  evaluated  for  establishing  safe  seismic  stress/strain  levels  for  the  piping  systems  of 
concern in the present research project, within the design framework of ASME B31.3 and EN 13480-3. 
This evaluation should be based on the various failure modes (mainly collapse and fatigue) and account 
for the influence of pipe diameter, schedule, D/t, material, material properties, hoop pressure stress, and 
fabrication  details.  Additional  testing  and  relevant  nonlinear  finite  element  parametric  studies  of  the 
component  tests  are  necessary  towards  this  effort,  and  will  be  conducted  in  the  framework  of  the 
present project. 

III.1.2.  Task 1.3: Basic seismic structural design of typical liquid storage tanks 

The present task focuses on the seismic design of liquid storage tanks. The main purposes of this study 
are  

  the identification of significant differences between the American and the European standards.  

  to pin-point inconsistencies and missing links in design 

  to identify issues that require further consideration  

77 
 
 
 
Figure III.1.2-1: Three liquid storage tanks which are used in the design examples 

For a detail description of the analysis and design of the structures the reader is referred to deliverable 
D1.3. 

III.1.3.  Task  1.4:  Basic  seismic  structural  design  of  typical  cylindrical  and  spherical 

pressure vessels 

The present task focuses on the seismic design of pressure vessels. The main purposes of this study are 

  outlining the current methodology for pressure vessel design with regard to seismic loading 

  summarizing the provisions of European and American standards for the treatment of earthquake 

loads 

  pin-pointing  important  design  issues  (e.g.  missing  links,  inconsistencies)  that  require  further 

investigation 

(a) 

(b) 

Figure III.1.3-1: Sketch of (a) vertical pressure vessel on skirt support and (b) horizontal pressure vessel on saddle 
supports (drawings made by software PV Elite 2010). 

For a detail description of the analysis and design of the structures the reader is referred to deliverable 
D1.4. 

III.1.4.  Task  1.5:  Basic  seismic  structural  design  of  typical  piping  system  and  its 

support structure 

The  main  goal  of  this  study  is  to  illustrate  in  detail  the  seismic  design  of  piping  systems,  typical  of 
petrochemical refineries. Both the available American and European standards have been applied to an 
existing piping system, performing the prescribed verification steps. In particular, the following issues 
have been illustrated and discussed: 

  Calculation  of  seismic  forces  using  the  available  codes/specifications,  using  both  static  and 

dynamic analyses. 

78 
 
 
 
 
 
         
 
 
 
  Verification checks at some critical components 

  Comparison between European and American code results 

  Inconsistencies, issues that require further consideration; to be used in the following WPs. 

Figure III.1.4-1: The analyzed piping system on a pipe-rack 

Figure  III.1.4-2:  Configuration  of  10”  ammonia 
transmission line. 

Figure  III.1.4-3:  Configuration  of  20”  gas  transmission 
pipeline. 

III.2.  Appendix of WP2 

III.2.1.  Task 2.1: Seismic forces in vertical cylindrical liquid storage tanks 

Three  liquid  storage  tanks are existing  tanks,  constructed  in  seismic regions,  and  have  been  designed 
against  hydrostatic pressure  following  the  API  650  methodology  “Calculation of Thickness by  the  1-
Foot Method” in section 5.6.3, considering a liquid containment of unit gravity (water). The tanks are 
depicted in Figure III.2.1-1. 

Figure III.2.1-1:  Tank I (left), tank II (center) and tank III (right) used for the parametric studies in the present 
paper. 

   A B 1 2 3 4 5 6 7 x y z 79 
 
 
 
 
 
 
 
(a) 

(b) 

Figure III.2.1-2: Finite element analysis of unanchored liquid storage tank under lateral loading; (a) uplifting of 
base plate, (b) development of elephant’s foot at the tank bottom. 

Figure III.2.1-3: Elephant’s foot buckling, as obtained by finite element simulation. 

III.2.2.  Task 2.2: Seismic forces in pressure vessels 

In  this  task,  the  seismic  analysis  and  design  indicated  that  more  detailed  information  regarding  the 
estimation of the behavior factor for these structures should be included in the design codes, as well as 
more  detailed  guidelines  for  the  determination  of  the  seismic  forces.  An  attempt  to  estimate  the 
behavior factor for these structures is presented using advanced numerical tools, for pushover analysis 
and  incremental  dynamic  analysis.  The  pushover  analysis  (see  Figure  III.2.2-1  and  Figure  III.2.2-2) 
indicated that the capacity  of  the  presure  vessels is  quite  higher  than the  required  strength  in  seismic 
loading.  On  the  other  hand,  using  incremental  dynamic  analysis  for  the  vertical-cylindrical  pressure 

upliftingLwElephant’s foot formation80 
 
 
 
  
 
 
 
 
 
vessel, the value of behavior factor calculated is somewhat below the corresponding value suggested by 
American standard ASCE 7. 

Figure III.2.2-1: Von Mises stress distribution at maximum capacity (approximately 300mm displacement) 

(a) 

   (b)

Figure III.2.2-2: Von Mises stress distribution for the horizontal pressure vessel in the (a) longitudinal x-direction 
(approx. 10 mm displacement) and (b) transverse z-direction  (approx. 50 mm displacement). 

III.3.  Appendix of WP3 

III.3.1. 

III.3.1.1.

Task 3.1 Basic material testing and characterization 

Base metal testing of X52 and P355 material extracted from elbows, pipes and plates 

Longitudinal and transversal specimens have been extracted from X52 elbows at flank position, where 
large  plastic  strain  accumulation  is  present  during  bending.  Full  thickness  strip  specimens  have  also 
been extracted in longitudinal direction from X52 pipes to be connected to elbows in full scale testing 
The  results  are  summarized  in  Table  III.3.1-1  and  Table  III.3.1-2  and  shows  that  actual  material 
properties fulfil the requirements of ISO3183 standard [33], for X52 steel elbows and EN 10261-3 for 
P355  steel  grade  pipes  [21].  It  can  be  observed  that  very  little  difference  exists  between  these  two 
materials. 

81 
 
 
 
 
 
 
 
 
Table III.3.1-1: Monotonic tensile properties of API X52 specimen extracted from elbow 

Pipe type 

Elbow 

Specimen id. 

E2B-L10  E2B-L11  E2B-T1  E2B-T2 

Direction 

Longitudinal 

Transversal 

Rp0.2% 

Yield Strength (MPa) 

R m 

A% 

Tensile Strength (MPa) 

Elongation (%) 

412 

569 

40 

425 

577 

39 

400 

580 

41 

399 

578 

39 

Table III.3.1-2: Monotonic tensile properties of API5L X52  specimens extracted from pipes to be used in elbow 
piping specimens 

Pipe type 

X52 Pipe 

Specimen id. 

ASTM Strip-1 

ASTM Strip-2 

Direction 

Longitudinal 

Rp0.2% 

Yield Strength (MPa) 

R m 

A% 

Tensile Strength (MPa) 

Elongation (%) 

397 

543 

36 

396 

537 

31 

Table III.3.1-3: Summary of monotonic tensile properties of P355N extracted from straight pipes 

Pipe type 

P355N 

Specimen id. 

ASTM Strip-1 

ASTM Strip-2 

Direction 

Longitudinal 

Rp0.2% 

Yield Strength (MPa) 

R m 

A% 

Tensile Strength (MPa) 

Elongation (%) 

387 

528 

40 

388 

557 

40 

Cyclic  testing  on  X52  extracted  from  elbow,  reported  in  Figure  III.3.1-1  and  Figure  III.3.1-2,  shown 
that little hardening or softening can be reported for this material, even if some softening seems present 
and  its  level  reduces  when  increasing  the  strain  cyclic  range,  Figure  III.3.1-1.  Ratcheting  testing, 
reported  in  Figure  III.3.1-2,  revealed  that  some  difference  in  the  accumulation  of  the  strain  can  be 
reported between longitudinal and transversal direction being the longitudinal direction more prone to 
accumulation  of  strains.  Furthermore  it  can  be  reported  that  the  evolution  of  strain  accumulation  is 
linear for the transversal direction while is always increasing in the longitudinal direction. 

Figure  III.3.1-1:  API  X52  from  elbows  cyclic  properties  at  various  imposed  strain  levels,  longitudinal  direction 
(left) transverse direction (right) 

-600-400-2000200400600-2.5-1.5-0.50.51.52.5Stress -MPaStrain -%E1A-L-3 strain +/-2.25%E1A-L-2 strain +/-1.5%E1A-L-4  strain +/-0.75%-600-400-2000200400600-2.5-1.5-0.50.51.52.5Stress -MPaStrain -%E1B-T-5 strain +/-2.25%E1B-T-9 strain +/-1.5%E1B-T-4   strain +/-0.75%82 
 
 
 
 
 
 
Figure  III.3.1-2:  API  X52  from  elbows  cyclic  properties  at  various  imposed  stress  levels,  longitudinal  direction 
(left) transverse direction (right) 

Finally, results of small scale testing on P355 specimen extracted from plates to used in the nozzles full 
scale testing, are reported in the following Table III.3.1-4, for monotonic testing and in Figure III.3.1-3, 
for cyclic testing at various levels of imposed strain. Again little hardening or softening is observed in 
strain controlled cyclic tests on plates, thus confirming that P355 and X52 material properties show The 
measured  cyclic  properties  have  been  used  by  RWTH  to  calibrate  the  nonlinear  kinematic  hardening 
model employed in FE simulations of nozzle components. 

Table III.3.1-4: Summary of monotonic tensile properties of P355 extracted from plates 

Material 

P355 

Specimen id. 

ASTM Strip-1 

ASTM Strip-2 

Direction-1 

Rp0.2% 

Yield Strength (MPa) 

R m 

A% 

Tensile Strength (MPa) 

Elongation (%) 

Direction-2 

Rp0.2% 

Yield Strength (MPa) 

R m 

A% 

Tensile Strength (MPa) 

Elongation (%) 

366 

534 

46 

369 

529 

47 

362 

530 

46 

397 

532 

46 

Figure III.3.1-3: P355 from elbows cyclic properties at various imposed stress levels 

02468101214160100200300400Strain -%Cycle numberE2A-L-7 -Stress 3 R = -0.8E2A-L-6 -Stress 2 R = -0.6E2A-L-9 -Stress 1 R = -0.402468101214160100200300400Strain -%Cycle numberE1B-T-8 -Stress 3 R = -0.8E1B-T-7 Stress 2 R = -0.6E1B-T-6 -Stress 1 R = -0.4-600-400-2000200400600-0.03-0.02-0.0100.010.020.03Stress -[MPa]strain [-] P355 plate Δε= +/-2.25%P355 plate Δε= +/-1.5%P355 plate Δε=+/-0.75%83 
 
 
 
 
 
 
III.3.1.2.

Base metal testing of material to be used for the tee-joint specimens 

Monotonic tensile tests on coupon specimens, extracted from the 8” and 6” pipes and from the tee-joint 
8”x6” fitting to be used in WP3 Task 3.5, were carried out to define the true stress-strain curve of the 
material. The specimens were taken along the length of the pipes and along the 8” straight part of the 
fitting.  The  tensile  tests  and  geometry  selection  are  in  accordance  with  EN  ISO  6892.01  [34].  The 
material  standard  specifies  a  minimum  yield  point  of  355  MPa,  ultimum  strength  of  490  MPa  and 
elongation  22%  for  the  P355NH/NL1  material.  The  data,  presented  in  the  table  below,  are  in 
agreements with both the EN standard and the material certificates.  

Table III.3.1-5: Summary of monotonic tensile properties of the base material. 

8” pipe 

6” pipe 

8”x6” fitting 

Yield strength (MPa) 

Tensile strength (MPa) 

Re 

Rm 

410 

541 

414 

542 

418 

549 

409 

538 

410 

540 

412 

543 

378 

571 

369 

548 

Elongation (%) 

A 

23.3 

23.4 

26.3 

26 

26.1 

26.4 

30.4 

27.1 

III.3.1.3.

Testing of welded specimens from P355NH/NL1 material  

Butt  welded  specimens  were  manufactured from  8”  pipe  supplied  for  the  needs of  Tee-joint  medium 
scale tests in Task 3.5. The WPS No 02 (given in D2.3) was used. The test program defined for welding 
certification (EN15614.01 [27]) was performed in order to check the validity of the WPS for butt welds 
that will be used in the medium scale tests. 

Firstly,  non-destructive  tests  were  carried.  No  defects  were  detected  by  magnetic  particle  and 
radiographic  inspection.  Destructive  tests  were  also  performed.  Namely,  macroscopic  examination, 
hardness  measurement,  bend,  transverse  tensile  and  impact  tests  were  carried  out.  Overall,  the  test 
results for the certification of the butt-welded specimens using the TIG method were satisfactory. 

III.3.1.4.

Cyclic testing of welded specimens from P355NH/NL1 material  

Uni-axial cyclic tests of welded specimens have been carried out at varying stress amplitudes. First a 
monotonic test was carried out to confirm the yield point of the joint which was measured to be 400 
MPa. Then 6 coupon specimens were tested at different levels of amplitude and a mean stress value of 
σm=Re/2=200 MPa. The stress range was gradually increased at specified multiples of the yield stress 
Δσ=a%  x  Re.  The  stress  amplitude  versus  the  number  of  cycles  until  failure  is  plotted  in  Figure 
III.3.1-4.  Two  more  specimens  were  tested  at  different  levels  of  mean  stress  and  amplitude  and  the 
results are also plotted in Figure III.3.1-4. The first was tested from σmin = 0 MPa to σmax = 464 MPa (σm 
= 232 MPa, Δσ = 464 MPa) and the second from σmin = -410 MPa to σmax = 410 MPa (σm = 0, Δσ = 810 
MPa).  

Figure III.3.1-4: Stress amplitude versus number of cycles until failure. 

0200400600800010.00020.00030.00040.00050.00060.00070.00080.000Number of cyclesStress amplitude Δσ (MPa)σm=200MPaσm=0σm=232MPaΛογαριθμική (σm=200MPa)84 
 
 
 
 
 
 
 
 
 
 
In all specimens apart from the last one, with the largest compressive stress applied, cracks initiated at 
the  HAZ  near the  weld root.  In  some  specimens  cracks  also started from  the weld  cap.  When  cracks 
initiated  from  the  weld  root  the  always  propagated  through  the  weld  metal.  In  two  specimens  cracks 
developed  in  the  HAZ  of  both  the  weld  root  and  cap.  In  those  the  crack  that  started  from  the  cap 
propagated quicker through the base metal of the thinner part rather than the ones that initiated from the 
root and propagated though the weld.  For a mean value of stress equal to 0 and stress amplitude 820 
MPa the mode of failure was quite different. The crack initiated in the HAZ area near the weld cap and 
quickly propagated through the base metal of the thinner part of the joint. Failure occurred at only 880 
cycles. 

III.3.2. 

III.3.2.1.

Task 3.2: Tests on tank openings/nozzles under strong cyclic loading  

Numerical results for monotonic loading 

For  all  3  types  of  shell  nozzle  reinforcements  4  different  loading  conditions  were  investigated:  (1) 
tension, (2) compression, (3) shear upward and (4) shear downward acting onto the nozzle (see  Figure 
III.3.2-1), where hereafter tension and compression are also referred to longitudinal load direction (-1) 
and the shear loadings are referred to transverse load direction (-2). They were used the determination 
of the cyclic test procedure. 

Figure III.3.2-1: Load directions for monotonic loading 

Table III.3.2-1 summarizes the yield loads Fy and the corresponding displacement ey obtained from the 
monotonic load-displacement curves regarding the ECCS-recommendation [20]. 

Table  III.3.2-1:  Yield  loads  and  corresponding  displacements  of  the  internal-,  external-  and  thickened  nozzle 
reinforcement for different load directions 
1 Tension 

2 Compression 

4 Shear down 

3 Shear up 

Specimen 

P1 Internal 

P2 External 

P3 Nozzle 

[kN] 
677.2 
583.1 
455.0 

[mm] 
9.9 
9.9 
8.1 

[kN] 
481.1 
427.8 
494.4 

[mm] 
8.1 
8.6 
10.6 

[kN] 
178.5 
148.2 
101.1 

[mm] 
24.5 
23.1 
18.2 

[kN] 
178.8 
146.4 
101.1 

[mm] 
24.2 
22.5 
18.1 

III.3.2.2.

Test results for cyclic loading 

The test results in terms of cyclic load-displacement curves referring to the point of load application and 
pictures  of  the  specimens  at  the  end  of  the  loading  history  are  shown  in  Figure  III.3.2-2  and  Figure 
III.3.2-3.  Due  to  slippage  in  the  bolts  of  the  load  application  the  test  specimens  were  loaded  with 
slightly smaller displacements than indicated by the true movement of the jack. 

yFyeyFyeyFyeyFye85 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
Figure III.3.2-2: P1-1, P2-1 and P3-1– Cyclic load-displacement curve and pictures of the specimens at the end of 
the loading history (longitudinal load direction) 

-600-400-200020040060080010001200-100-80-60-40-20020406080100Load [kN]Load-point displacement [mm]cycle 1-4cycle 5-7cycle 8-10cycle 11-20cycle 21-23cycle 24-26cycle 27-38P1-1 Internal-longitudinal-600-400-200020040060080010001200-100-80-60-40-20020406080100Load [kN]Load-point displacement [mm]cycle 1-4cycle 5-7cycle 8-10cycle 11-20cycle 21-23cycle 24-26cycle 27-38P2-1 External-longitudinal-600-400-200020040060080010001200-100-80-60-40-20020406080100Load [kN]Load-point displacement [mm]cycle 1-4cycle 5-7cycle 8-10cycle 11-20cycle 21-23cycle 24-26cycle 27-30P3-1 Nozzle-longitudinal86 
 
 
 
 
Figure III.3.2-3: P1-2, P2-2 and P3-2– Cyclic load-displacement curve and pictures of the specimens at the end of 
the loading history (transverse load direction) 

-300-200-1000100200300-100-80-60-40-20020406080100Load [kN]Load-point displacement [mm]cycle 1-4cycle 5-7cycle 8cycle 9-11cycle 12-13P1-2 Internal-transverse-300-200-1000100200300-100-80-60-40-20020406080100Load [kN]Load-point displacement [mm]cycle 1-4cycle 5-7cycle 8cycle 9-12cycle 13-15P2-2 External-transverse-300-200-1000100200300-100-80-60-40-20020406080100Load [kN]Load-point displacement [mm]cycle 1-4cycle 5-7cycle 8-10cycle 11-17cycle 18-27P3-2 Nozzle-transverse87 
 
 
 
 
III.3.3.  Task 3.3: Tests on bolted piping connections under strong cyclic loading 

III.3.3.1.

Tests on standard bolted piping connections under strong cyclic loading 

III.3.3.1.1 

Test instrumentation 

Full  scale  testing  instrumentation  has  been  employed  to  measure,  monitor  and  control  the  main  test 
parameters:  forces,  displacements,  local  strains  and  pressure.  Measured  data  have  been  employed  in 
order to numerically simulate (by means of FEA) the experimental tests and to provide a tuning of the 
numerical  tools.  A  plan  of  test  instrumentation  positioning  is  reported  in  Figure  III.3.3-1  and  Figure 
III.3.3-2. It included: 

  Load cell to control and measure the applied bending load;  

  Strain gauges to measure strains on the pipe at critical locations, Figure III.3.3-1; 

  LVDT: to control and measure applied displacements and rotations, Figure III.3.3-2; 

  Pressure transducers to monitor and control the internal pressure level.  

A dedicated Digital acquisition system has been developed on purpose by CSM. 

Figure III.3.3-1: Strain gauges position mapping  

Figure III.3.3-2: Measured deflection at mid sections and rotation at load application points 

III.3.3.1.2 

Test results 

Test  results  in  terms  of  actuator’s  load  vs.  stroke  diagrams  pictures  of  the  specimen  after  testing  are 
reported in the following. 

88 
 
 
 
 
 
 
 
Figure  III.3.3-3:  Actuator  load  vs.  displacement  Test 
PN63-1 

Figure III.3.3-4: PN63-1 specimen after testing 

Figure  III.3.3-5:  Actuator  load  vs.  displacement  Test 
PN40-1 

Figure III.3.3-6: PN40-1 specimen after testing 

Figure  III.3.3-7:  Actuator  load  vs.  displacement  Test 
PN63-2 

Figure III.3.3-8: PN63-2 specimen after testing 

020406080100120140160180200050100150200250300350400450050100150Load [kN]Stroke [mm]LoadDeflectionStart of leakeage 371kNDeflection [mm]buckling-800-600-400-2000200400600800-60-40-200204060Load [kN]Stroke [mm]89 
 
 
 
 
 
 
 
Figure  III.3.3-9:  Actuator  load  vs.  displacement  Test 
PN63-3 

Figure III.3.3-10: PN63-3 specimen after testing 

Figure  III.3.3-11:  Actuator  load  vs.  displacement  Test 
PN40-2 

Figure III.3.3-12: PN40-2 specimen after testing 

Figure  III.3.3-13:  Actuator  load  vs.  displacement  Test 
PN40-3 

Figure III.3.3-14: PN40-3 specimen after testing 

Figure  III.3.3-15:  Actuator  load  vs.  displacement  Test 
PN40-4 

Figure III.3.3-16: PN40-4specimen after testing 

-600-400-2000200400600-30-20-100102030Load [kN]Stroke [mm]-500-400-300-200-1000100200300400500-60-40-200204060Load [kN]Stroke [mm]-400-2000200400-30-20-100102030Load [kN]Stroke [mm]-500-400-300-200-1000100200300400500-60-40-200204060Load [kN]Stroke [mm]90 
 
 
 
 
 
 
 
 
III.3.3.2.

Tests on non-standard bolted piping connections under strong cyclic loading 

Figure III.3.3-17: (a) Designed non-standard BFJs and dimensions; (b) spiral wound gasket; (c) stud bolt 

Table III.3.3-1: Dimensions of the designed non-standard BFJs (mm) 

Pipe 
Size 

O  W  G  K 

T 

J 

A 

B  H 

P  Q  M  N 

No 
of 
Bolts 

Stud 
Bolt 
Size 

DN 200 
5
7
SCH 40  3

0
2
3

0
9
2

0
3

18 (Design 
01) 
27 (Design 
02) 

8
1
.
8

5
.
1
2
2

4
7
.
2
0
2

1
.
9
1
2

6
1
2

8
2
2

8
4
2

0
9
2

12 

M 27 x 
3.00 

Table III.3.3-2:  Test program on non-standard BFJ carried out by UTrento 

No. 

Test type 

Test name 

Loading type 

Internal pressure 

Specimen 

1 
2 
3 
4 
5 
6 
7 
8 

Bending 
Bending 
Bending 
Bending 
Axial 
Axial 
Axial 
Axial 

BSML18 
BSML27 
BSCL18 
BSCL27 
ASML18 
ASCL18 
ASCL27-1 
ASCL27-2 

Monotonic 
Monotonic 
Cyclic 
Cyclic 
Monotonic 
Cyclic 
Cyclic 
Cyclic 

1.5 MPa 
1.5 MPa 
1.5 MPa 
1.5 MPa 
1.5 MPa 
1.5 MPa 
1.5 MPa 
1.5 MPa 

BS1 
BS2 
BS3 
BS4 
AS1 
AS2 
AS3 
AS4 

Figure III.3.3-18: (a) Bending test specimen, BS1; (b) axial test specimen, AS1; (c) a spiral wound gasket; (d) a 
stud bolt. 

91 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
(a) 

(b) 

Figure III.3.3-19: Test set-up for (a) bending tests; (b) axial tests. 

(a) 

(c) 

(b) 

(d) 

Figure  III.3.3-20:  (a)  Actuator  load-displacement  curve  from  test-  BSML18;  moment-rotation  curves  from  (b) 
test- BSCL18, (c) test- BSML27 and (d) test- BSCL27. 

050100150200250050100150200250300350400450Displacement, mmLoad, kN  BSML18-80-60-40-20020406080100-200-150-100-50050100150200Rotation, milliradMoment, kNm  BSCL18PushPull01020304050607080050100150200250Rotation, milliradMoment, kNm  BSML27-40-200204060-200-150-100-50050100150200Rotation, milliradMoment, kNm  BSCL27PullPush92 
 
 
 
 
 
 
 
 
 
(a) 

(c) 

(b) 

(d) 

Figure  III.3.3-21:  Load-flange  displacement  curve  from  (a)  test  ASML18,  (b)  test  ASCL18,  (c)  test  ASCL27-1 
and (d) test ASCL18-2. 

Figure III.3.3-22: Failure of the pipe and bending of a bolt during test BSCL27 

0123456780200400600800100012001400160018002000Displacement, mmLoad, kN  ASML180123456-1000-5000500100015002000Displacement, mmLoad, kN  ASCL1800.511.52-1000-5000500100015002000Displacement, mmLoad, kN  ASCL27-100.511.52-5000500100015002000Displacement, mmLoad, kN  ASCL27-293 
 
 
 
 
 
 
 
 
 
 
Table III.3.3-3: Comparison between experimental leakage moment and allowable moments suggested by codes 

Experimental moment 

Allowable moments by codes 

Minimum leakage moment obtained 
from bending tests 

EN 13480, 2002  ASME B31.1 & B31.3 

80.24 kNm 

51.23 kNm 

57.08 kNm 

Table III.3.3-4: Comparison between experimental leakage load and allowable loads suggested by codes 

Experimental load 

Allowable loads by codes 

Minimum leakage load obtained  
from axial tests 

EN 13480, 2002 

ASME B31.1 & B31.3 

1170 kN 

885.20 kN 

885.20 kN 

Table III.3.3-5: Maximum moment, Mmax, axial force, Nmax and shear force, τmax in the piping system obtained 
from a case study 

Mmax 

16.79 kNm 

Nmax 

180.5 kN 

τmax 

5.08 N 

III.3.4. 

Task 3.4: Tests on piping elbows under strong cyclic loading 

III.3.4.1.

Tests on non-pressurized piping elbows under strong cyclic loading 

The evolution of the load-displacement cycles for each one of the eight tests conducted is presented in 
detail in Figure III.3.4-4 to Figure III.3.4-11. For the lower end-displacement amplitudes (Test 1 to Test 
4), the resulting load-displacement loops are not changing in shape except for the cycle where failure 
takes  place.  On  the  contrary,  in  Tests  5,  6  and  7,  the  initial  shape  and  size  of  the  load-displacement 
cycles  changes  significantly  indicating  a  loss  of  resistance  (structural  degradation),  as  the  number  of 
cycles increases. This phenomenon can be attributed to the accumulation of plastic deformation at the 
most  strained  locations  of  the  elbow  resulting  to  the  formation  of  plastic  hinges  at  the  elbow  middle 
section and finally to cracking of the elbow, see Figure III.3.4-2. 

Cross-sectional “flattening” is measured at two positions along the specimen. The first position is at the 
middle section of the elbow, where maximum deformation occurs, while the second location is near the 
weld region, as shown in Figure III.3.4-2. Two representative flattening-displacement curves are shown 
in Figure III.3.4-8 and Figure III.3.4-12.  

Figure  III.3.4-1:  Cracked  elbow  specimen  No.  4  with 
displacement  range  ±150  mm:  Failure  stage  with  through 
thickness crack development  

Figure 
Instrumentation 

III.3.4-2: 

Experimental 

set-up: 

94 
 
 
 
 
 
 
 
 
 
 
Figure  III.3.4-3:  Increasing  amplitude  –  Loading  pattern 
according to ECCS recommendations. 

Figure  III.3.4-4:  Test  1  - 
displacement cycles  

    Load  – 

Figure  III.3.4-5:  Test  2  - 
displacement cycles 

  Load  – 

Figure  III.3.4-6:  Test  3  - 
displacement cycles 

  Load  – 

Figure  III.3.4-7:  Test  4  - 
displacement cycles  

    Load  – 

Figure III.3.4-8: Test 4 - 
– displacement cycles 

 Flattening 

25 lmm70 lmm100 lmm150 lmm150 lmm95 
 
 
 
 
 
 
 
 
 
 
Figure  III.3.4-9:  Test  5  -   
displacement 

  Load  – 

Figure  III.3.4-10:  Test  6  - 
displacement 

  Load  – 

Figure  III.3.4-11:  Test  7  - 
displacement cycles  

  Load  – 

Figure 
Flattening – displacement cycles 

III.3.4-12:  Test  7 

- 

III.3.4.2.

Tests on pressurized piping elbows under strong cyclic loading 

CSM Full Scale Test Machine (FSTM) Figure III.3.4-13, dedicated to the testing of tubular components 
has  been  used for elbow  testing,  some  modifications  have  been  made  in  order to  perform  the  test,  in 
particular the specimen is connected to the machine by  means of specially designed hinged clamping 
system that allows for rotations of both ends, while one end is horizontally displaced by actuator. Main 
machine capabilities in terms of loads and displacements are reported in Table III.3.4-1. 

Load 
capacity 

Actuator 
displacement  

2500 tons, tensile  
 3000, 
compressive 

+/- 350mm  
700 
direction 

in 

one 

: 

: 

Internal 
pressure 

:  1200 bar 

Figure  III.3.4-13:  CSM  FSTM  in  full  scale  testing 
machine 

Table III.3.4-1:FSTM main capabilities 

200 lmm250 lmm300 lmm300 lmmFrameTension and Compression loadClamping systemBending jacks96 
 
 
 
 
 
 
 
 
 
 
 
 
The test the following parameters of the test have been monitored: 

elbow displacements by LVDT’s as for Figure III.3.4-14; 

 
  Mid-section vertical and horizontal ovalization, by means of the equipment of Figure III.3.4-15; 
  Strains, Figure III.3.4-16. 
  Applied force by means of a load cell. 

Figure  III.3.4-17  to  Figure  III.3.4-21  show  the  load displacement  curves  for  all  tests  and the through 
wall  crack  on  elbows.  Figure  III.3.4-22  and  Figure  III.3.4-23  show,  as  an  example,  the  ovalization 
diagrams for tests 2 and 4 with  the same pressure (70 bar). Whole test data is reported in deliverable 
D3.4. 

Figure III.3.4-14: Measure of elbow displacements 

Figure III.3.4-15: Measure of ovalization 

Figure III.3.4-16: Strain gauges installation 

Figure III.3.4-17: Test 1 load displacement curve (left) and final crack position (right) 

LVDT 1  LVDT 2  -120-80-4004080-220-160-100-402080140200Force -kNDisplacement -mmTest-1 CSM +/-200mm p=32bar97 
 
 
 
 
 
 
 
 
 
Figure III.3.4-18: Test 2 load displacement curve (left) and final crack position (right) 

Figure III.3.4-19: Test 3 load displacement curve (left) and final crack position (right) 

Figure III.3.4-20: Test 4 load displacement curve (left) and final crack position (right) 

-120-80-4004080-220-160-100-402080140200Force -kNDisplacement -mmTest-2 CSM +/-200mm p=70bar-160-120-80-4004080-320-240-160-80080160240320Force -kNDisplacement -mmTest-3 CSM +/-300mm p=32bar-160-120-80-4004080-320-240-160-80080160240320Force -kNDisplacement -mmTest-4 CSM +/-300mm p=70bar98 
 
 
 
 
 
 
 
 
 
Figure III.3.4-21: Test 5 load displacement curve (left) and final crack position (right) 

Figure III.3.4-22: Test 2 elbow ovalization  

Figure III.3.4-23: Test 4 elbow ovalization 

-120-80-4004080120-220-160-100-402080140200Force -kNDisplacement -mmTest-5 CSM +/-200mm p=120bar180190200210220230240250260270-320-240-160-80080160240320Diameter -mmApplied displacement -mmOD vert. actual Test-2 CSM +/-200mm p=70barOD horiz actual Test-2 CSM +/-200mm p=70bar180190200210220230240250260270-320-240-160-80080160240320Diameter -mmApplied displacement -mmOD vert. actual Test-4 CSM +/-300mm p=70barOD horiz actual Test-4 CSM +/-300mm p=70bar99 
 
 
 
 
 
 
 
 
III.3.4.3.

Pseudo-dynamic loading on piping system 

Table III.3.4-2: Experimental program on a piping system carried out by UTrento 

Test Case 

Excitation 
Wave 

Excitation 
Direction 

PGA 
m/s2 

PGA 
g 

Identification 
tests 

Identification test of the 
physical substructure, 
IDT 

Hammer  
Test 

Hammer 
impact 

- 

- 

- 

Elastic tests 

Elastic tests, ET 

PDDS  

Seismic 

Horizontal  0.4128 

0.0421 

Serviceability 
limit state 
tests 

Ultimate 
limit state 
tests 

Operational limit state 
test, SLOT 
Damage limit state test 
SLDT 
Safe life limit state test 
SLVT 
Collapse limit state tests 
SLCT 

PDDS 

Seismic 

Horizontal  0.7719 

0.0787 

PDDS 

Seismic 

Horizontal  1.0980 

0.1119 

PDDS 

Seismic 

Horizontal  4.1278 

0.4208 

PDDS 

Seismic 

Horizontal  5.8779  0.59918 

Table III.3.4-3: PGA corresponding to different limit states 

Limit States 

PGA, m/sec2  

PGA, Ratio 

Serviceability 

SLO 

Operational limit state 

0.77 

0.187 SLV 

Limit States 

SLD 

Damage limit state 

Ultimate 

SLV 

Safe life limit state  

Limit States 

SLC 

Collapse limit state 

1.1 

4.13 

5.88 

0.266 SLV 

1.000 SLV 

1.424 SLV 

Table  III.3.4-4:  Maximum  moment,  Mmax  stress,  σmax  and  shear,  τmax  in  the  piping  system  under  reference 
seismic loading 

Response without damping 

Component 
Elbow 2 
Straight pipe 

Mmax (kNm) 
69.15 
81.56 

σmax (MPa) 
503.56 
296.03 

Stress level 
Above yield strength 
Below yield strength 

Response with Rayleigh damping 

Component 
Elbow 2 
Straight pipe 

Mmax (kNm) 
43.63 
45.62 

σmax (MPa) 
317.74 
288.12 

Stress level 
Below yield strength 
Below yield strength 

τmax in the straight pipe (kN) 

τmax (without damping) 
τmax (with damping) 

95.48 
48.50 

100 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
Figure III.3.4-24: (a) A 3D model of the piping system and support structure; (b) PS, NS and coupling nodes 

(a)   

(b) 

Figure III.3.4-25: Schematic of the experimental set-up for the hybrid tests 

Figure III.3.4-26: Actual test set-up for the PDDS 

101 
 
 
 
 
 
 
 
 
An elbow 

The tee joint 

The bolted flange joint 

Figure III.3.4-27: Some components of the test specimen 

(a) 

(b) 

Figure III.3.4-28: Rotation histories and relevant spectra of the two end nodes of Elbow #1 at (a) SLVT and (b) 
SLCT 

Figure III.3.4-29: Strain histories of Elbow #2 at (a) SLVT and (b) SLCT 

(a)  

(b)  

102 
 
 
 
 
 
 
 
 
 
(a) 

(b) 

Figure III.3.4-30: Displacement histories and relevant spectra of the Coupling Node #1 at (a) SLVT and (b) SLCT 
(Blue: Experimental; Magenta: Numerical) 

III.3.5.  Task 3.5: Tests on T-joint pipe connections under strong cyclic loading 

For  the  experimental  study  of  branch-to-main-pipe  T-joint  connections  under  strong  fully-reversing 
cyclic  loading,  two  (2)  connection  types  were  considered,  a  connection  with  a  fitting  (TJF)  and  a 
connection with a welded reinforcement ring-plate (TJWR). In total, eight (8) specimens were tested, 
one (1) under static monotonic and three (3) under fully reversing cyclic loading for each connection 
configuration. 

To  conduct  these tests, the  main  steel  pipe  of  the  specimen  was  fixed  at  both  of  its  ends  to  the  steel 
reaction  testing  frame.  Both  ends  of  the  main  pipe  were  bolted  to  steel  columns  at  a  large  enough 
distance from the T-joint so that the fixity at the supports would not affect the behavior at the critical 
region of the connection. The load was applied through a hydraulic actuator hinged to the free-end of 
the branch through a steel ball-joint hinge. The applied load was perpendicular to the pipe-branch plane. 
Regarding the instrumentation setup for both types of specimens, two (2) wire position transducers were 
used to measure the load-point displacement (LPD) and a number of strain gages were placed at critical 
locations of the T-joint connection. In particular, axial 5-element strip strain gages were attached to the 
specimen in the fitting and welded reinforcement plate region for determining the strain distribution in 
the T-joint connection along the axis of the branch. The strain gages were at a distance of 3 mm, with 
the first one being placed 5 mm and 5078 mm away from the weld toe of the connection for the TJWR 
and TJF specimens, respectively.  

Two (2) of the T-joint specimens were subjected to static monotonic loading, one for each connection 
type.  Under  monotonic  loading  and  for  an  imposed  LPD  of  200  mm  the  specimen  with  the  fitting 
showed  about  93%  higher  load  carrying  capacity  than  that  of  the  specimen  with  the  welded 
reinforcement plate (see Figure III.3.5-1).  

Also, six (6) specimens were subjected to low-cycle fatigue three (3) for each connection type, up to a 
through-thickness  crack  failure  of  the  T-joint  connection.  Two  (2)  specimens  were  subjected  to  an 
imposed  LPD  value  of  δ1=60  mm  and  δ2=45  mm  and  one  (1)  specimen  was  subjected  to  a  fully-
reversing  cyclic  loading  history  according  to  the  ECCS  No45  [20]  testing  procedure.  For  the  ECCS 
testing  procedure,  an  LPD  value  of  δy=20  mm  was  considered  corresponding  to  a  conventional  yield 
load level Py based on the results of the tests under monotonic loading according to the recommended 
method of ECCS No 45. 

For  the  imposed  LPD  value  of  δ1=60  mm,  the  TJF  specimen  exhibited  a  higher  fatigue  strength  of 
about  7  times  that  of  the  TJWR  one.  The  TJF  specimen  failed  after  Nf=42  fully-reversing  loading 
cycles,  while  the  TJWR  specimen  after  Nf=6  loading  cycles  (see  Figure  III.3.5-2).  For  the  imposed 
displacement  of  δ2=45  mm,  the  TJF  specimen  failed  after  Nf=135  fully-reversing  loading  cycles 
exhibiting a higher fatigue resistance of about 8 times that of the TJWR one, which failed after Nf=17 
loading cycles (see Figure III.3.5-3). For the low-cycle fatigue tests according to ECCS No45, the TJF 

103 
 
 
 
specimen resisted 36% more loading cycles than the TJWR one. The TJF and TJWR specimen failed 
after Nf=15 and 11 loading cycles, respectively (see Figure III.3.5-4).  

All the T-joint specimens failed due to a through-thickness crack. The TJWR specimens failed along 
the circumference of the branch-to-reinforcement plate weld-toe (see Figure III.3.5-5). The specimens 
examined under monotonic loading and under an imposed LPD of 60 mm, failed in the branch region 
of  the  connection  while  the  other  two  specimens  in  the  welded  reinforcement  plate  region  of  the 
connection. In the case of the TJF specimens, cracking initiated at the top or the bottom of the fitting 
region in a bandwidth of about 20 mm transversely to the hoop direction of the main pipe at a distance 
of about 70 mm away from the branch-to-fitting weld-toe (see Figure III.3.5-6).   

Figure III.3.5-1: Load vs. LPD curves for the T-joint specimens under monotonic loading. 

                   (a)            

             (b) 

Figure III.3.5-2: Load vs. LPD curves under cyclic loading at an imposed LPD value of δ1=±60 mm: (a) TJF, (b) 
TJWR specimen. 

020406080100120140160180020406080100120140160180200220240Load-point deflection, δ (mm)Applied load, Pmax (kN)TJF-1MTJWR-1M-180-160-140-120-100-80-60-40-20020406080100120140160180-80-60-40-20020406080Load-point deflection, δ (mm)Load, Pmax (kN)δ=-60 mmδ=60 mm-180-160-140-120-100-80-60-40-20020406080100120140160180-80-60-40-20020406080Load-point deflection, δ (mm)Load, Pmax (kN)δ=60mmδ=-60 mmδ=60 mm104 
 
              
 
 
 
 
 
               (a)            

       (b) 

Figure III.3.5-3: Load vs. LPD curves under cyclic loading at an imposed LPD value of δ2=±45 mm: (a) TJF, (b) 
TJWR specimen. 

               (a)            

      (b) 

Figure  III.3.5-4: Load vs.  LPD  curves  under cyclic loading according to the ECCS No45 testing procedure: (a) 
TJF, (b) TJWR specimen. 

                                    (a)            
Figure  III.3.5-5:  Through-thickness  crack  failure  of  TJWR  specimen  at  the  weld  toe  of:  (a)  branch  region,  (b) 
main pipe region at the end of the welded reinforcement plate. 

                                             (b) 

-180-160-140-120-100-80-60-40-20020406080100120140160180-80-60-40-20020406080Load-point deflection, δ (mm)Load, Pmax (kN)δ=60mmδ=-45 mmδ=45 mm-180-160-140-120-100-80-60-40-20020406080100120140160180-80-60-40-20020406080Load-point deflection, δ (mm)Load, Pmax (kN)δ=-45 mmδ=45 mm-180-160-140-120-100-80-60-40-20020406080100120140160180-200-150-100-50050100150200Load-point deflection, δ (mm)Load, Pmax (kN)-180-160-140-120-100-80-60-40-20020406080100120140160180-200-150-100-50050100150200Load-point deflection, δ (mm)Load, Pmax (kN)105 
         
 
 
 
 
              
 
 
 
 
 
                   
 
 
 
 
 
Figure  III.3.5-6:  Through-thickness  crack  failure  in  the  region  of  the  fitting  for  the  TJF  specimen  subjected  to 
monotonic load. 

III.4. 

Appendix of WP4 

III.4.1. 

Task 4.3 Parametric numerical study of critical component behaviour 

III.4.1.1.

Parametric  numerical  study  of  elbow  behavior  under  extreme  monotonic  and  cyclic 
conditions 

III.4.1.1.1.  In-plane bending 

For each case examined, moment-rotation curves are derived and the local strain ranges at the critical 
locations are monitored. Analysis results for the case of 8-inch SCH40 long radius elbow are reported in 
Figure III.4.1-1 to Figure III.4.1-4.  

Figure  III.4.1-1:  Numerical  parametric  analysis 
,
results  –  8-inch  SCH40 

radius 

long 

Figure  III.4.1-2:  Numerical  parametric  analysis 
,
results  –  8-inch  SCH40 

radius 

long 

Figure  III.4.1-3:  Numerical  parametric  analysis 
radius, 
results 

SCH40 

8-inch 

long 

– 

Figure  III.4.1-4:  Numerical  parametric  analysis 
radius, 
results 

SCH40 

8-inch 

long 

– 

Δr=±0.02 radΔr=±0.04 radΔr=±0.08 radΔr=±0.12 rad106 
 
               
 
 
 
 
 
 
 
 
 
 
 
Parametric analysis results for the local strain range are reported in Table III.4.1-1 to Table III.4.1-4 for 
the complete set of cases. The strain ranges at the critical location of the elbow are reported for each 
case showing that the increased elbow wall thickness results to increased fatigue life for the same end-
rotation range due to the reduction of the strain field range developing localy. In addition, higher 
ratio results to increased fatigue life for the same end-rotation amplitude and schedule. 

8-inch SCH20 

Long radius 

Rotation 

Amplitude (rad) 

NLKH 

TL  NLKH 

TL 

8-inch SCH40 

Long radius 

Rotation 

Amplitude (rad) 

NLKH 

TL  NLKH 

TL 

0.69 

0.76 

1651 

1263 

0.67 

0.68 

1838 

2848 

1.66 

1.70 

133 

123 

1.61 

1.58 

143 

207 

4.87 

3.18 

7.88 

4.47 

5 

1 

20 

7 

6.00 

2.95 

6.70 

3.12 

5 

2 

29 

24 

Table  III.4.1-1:  Parametric  analysis  results:  8-inch 
SCH20 long radius 

Table III.4.1-2: Parametric analysis results: 8-inch 
SCH40 long radius 

8-inch SCH80 

Long radius 

Rotation 

Amplitude (rad) 

NLKH 

TL  NLKH 

TL 

8-inch SCH40 

Short radius 

Rotation 

Amplitude (rad) 

NLKH 

TL  NLKH 

TL 

0.59 

0.61 

2630 

2399 

0.77 

0.81 

1215 

1712 

1.46 

1.56 

193 

3.80 

1.67 

6.99 

2.46 

12 

2 

157 

130 

42 

2.01 

1.79 

76 

141 

6.02 

3.42 

9.78 

6.42 

3 

0 

18 

2 

Table III.4.1-3: Parametric analysis results:  8-inch 
SCH80 long radius 

Table  III.4.1-4: Parametric analysis results: 8-inch 
SCH40 short radius 

III.4.1.1.2.  Out-of-plane bending 

For  each  case  examined,  moment-rotation  curves  where  derived  and  the  local  strain  ranges  at  the 
critical  locations  where  monitored.  For  the  case  of  8-inch  SCH40  short  radius  results  are  reported  in 
Figure III.4.1-5 and Figure III.4.1-6 for three levels of pressure (0MPa, 3.2 MPa and 7.0 MPa). For the 
case of 8-inch SCH40 long radius results are reported in Figure III.4.1-7 for 0MPa pressure. Parametric 
analysis results are reported in Table III.4.1-5 to Table III.4.1-7 for the complete set of cases. The local 
strain ranges at the critical location of the elbow are reported for each case. 

R/D%Δε fN±0.02±0.04±0.08±0.12%Δε fN±0.02±0.04±0.08±0.12%Δε fN±0.02±0.04±0.08±0.12%Δε fN±0.02±0.04±0.08±0.12107 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
Figure III.4.1-5: Moment versus end-rotation curve 
for  displacement  range  ±0.25  rad,  (for  pressure  0 
MPa, 3.2 MPa and 7.0 MPa). 

Figure III.4.1-6: Moment versus end-rotation curve 
for  displacement  range  ±0.4  rad,  (for  pressure  0 
MPa, 3.2 MPa and 7.0 MPa). 

8-inch SCH80 

8-inch SCH80 

8-inch SCH40 

8-inch SCH40 

Long radius 

Short radius 

Long radius 

Short radius 

 Rotation 

Amplitude 
(rad) 

 Rotation 

Amplitude 
(rad) 

0.49 

4424 

0.83 

7.83 

10.8 

966 

1 

0 

0.87 

1.45 

12.72 

19.3 

849 

195 

0 

0 

0.55 

3132 

0.55 

3164 

0.91 

3.29 

5.54 

745 

18 

4 

0.98 

3.41 

5.72 

595 

16 

3 

Table III.4.1-5: Parametric analysis results: 8-inch 
SCH20 long and short radius. 

Table III.4.1-6: Parametric analysis results: 8-inch 
SCH40 long and short radius. 

 Rotation 

Amplitude 
(rad) 

8-inch SCH20 

Long radius 

0.87 

1.66 

8.78 

14.41 

839 

132 

1 

0 

Table  III.4.1-7:  Parametric  analysis  results:  8-inch 
SCH20 long radius. 

Figure  III.4.1-7:  Moment  versus  end-rotation  curve 
for displacement ranges ±0.05, ±0.1, ±0.25 and ±0.4 
rad, (pressure 0 MPa). 

III.4.1.2.

Parametric numerical study of flanged joints 

III.4.1.2.1 

Parametric numerical study of standard flanged joints 

The  typical  deformed  shape  observed  in  the  present  FEA  study  is  reported  in  Figure  III.4.1-8.  Local 
buckling  of  the  pipe  in  the  form  of  inward  directed  bulging  is  reported.  Within  the  buckle  large 
compressive  plastic  strains  are  present.  Although  these  strains  are  the  largest  observed  in  the  whole 
model, it is not likely that they can cause crackling. Large tensile strains are observed in the neck of the 

%Δε fN%Δε fN±0.05±0.1±0.25±0.4%Δε fN%Δε fN±0.05±0.1±0.25±0.4%Δε fN±0.05±0.1±0.25±0.4108 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
flange and on the pipe extrados, which have to be carefully considered as they could lead to through 
wall cracks.  

The  moment  rotation  curves  obtained  as  a  result  of  the  parametric  FE  study  are  reported  in  Figure 
III.4.1-9, Figure III.4.1-10 and Figure III.4.1-11. The peak moments that can be sustained by the various 
connections and the corresponding pipe end rotations are tabulated in Table III.4.1-8.  

The effect of the load parameters on the joint and a comparison for various geometries can be observed 
in Figure III.4.1-12 and Figure III.4.1-13. 

Figure III.4.1-8: typical deformed shape and plastic strain map of the flanged joints after bending 

Figure  III.4.1-9:  Moment  vs.  rotation  curves  for 
monotonic bending of PN40 flanged connections 

Figure  III.4.1-10:  Moment  vs.  rotation  curves  for 
monotonic bending of PN63 flanged connections 

Large tensile strainsrisk of crackingInward directed bulginglarge compressive strains 0204060801001201401600100200300400500End moment  -[kN*m]End  rotation -[millirad]PN40, p=0.5*PN, N=0.24*NyPN40, p=0.5*PN, N=-0.24*NyPN40, p=0.5*PN, N=-0.12*NyPN40, p=0.5*PN, N=0.12*NyPN40, p=1*PN, N=0PN40, p=0.5*PN, N=00204060801001201401601802000100200300400500End moment  -[kN*m]End  rotation -[millirad]PN63, p=0.5*PN, N=-0.24*NyPN63, p=0.5*PN, N=-0.12*NyPN63, p=0.5*PN, N=0.24*NyPN63, p=0.5*PN, N=0.12*NyPN63, p=1*PN, N=0PN63, p=0.5*PN, N=0109 
 
 
 
 
 
 
 
Figure III.4.1-11: Moment vs. rotation curves for monotonic bending of PN100 flanged connections 

Table III.4.1-8: Summary of buckling moment and rotation for monotonic behavior. 

Internal pressure effect, with N = 0 

PN40  

PN63  

PN100 

P/PN 

end rotation 

end moment  

end rotation 

end moment  

end rotation 

end moment  

- 

0 

0.5 

1 

millirad 

kNxm 

millirad 

kNxm 

millirad 

kNxm 

137 

165 

196 

139 

141 

142 

160 

189 

228 

178 

180 

180 

210 

230 

258 

270 

271 

271 

Axial load effect, with P = 0.5*PN 

PN40  

PN63  

PN100 

N/Ny 

end rotation 

end moment  

end rotation 

end moment  

end rotation 

end moment  

- 

millirad 

kNxm 

millirad 

kNxm 

millirad 

kNxm 

-0.24 

-0.12 

0 

+0.12 

+0.24 

100 

130 

165 

195 

195 

122 

132 

141 

146 

145 

110 

140 

189 

195 

175 

153 

167 

189 

185 

175 

115 

150 

230 

270 

251 

229 

250 

271 

280 

277 

Figure  III.4.1-12:  Effect  of  pressure  for  all  flange 
geometries 

Figure  III.4.1-13:  Effect  of  axial  load  for  all  flange 
geometries 

0501001502002503000100200300400500End moment  -[kN*m]End  rotation -[millirad]PN100, p=0.5*PN, N=-0.24*NyPN100, p=0.5*PN, N=-0.12*NyPN100, p=0.5*PN, N=0.12*NyPN100, p=100, N=0PN100, p=0.5*PN, N=0.24*NyPN100, p=50, N=00501001502002503000501001502002503000%20%40%60%80%100%End  moment -[kNxm]End  rotation -[millirad]Pressure -[% of PN]PN100 end rotationPN63  end rotationPN40  end rotationPN100 end moment PN63  end moment PN40  end moment 050100150200250300050100150200250300-24%-18%-12%-6%0%6%12%18%24%End  moment -[kNxm]End  rotation -[millirad]Axial load -[% of  Ny]PN40  end rotationPN63  end rotationPN100 end rotationPN40  end moment PN63  end moment PN100 end moment 110 
 
 
  
  
 
 
XIII.4.1.3.2 

Parametric numerical study of non-standard flanged joints 

(a) 

(b) 

Figure  III.4.1-14:  Stress  distribution  for  the  bending  test:  (a)  18  mm  thickness  joint,  and  (b)  36  mm 
thickness joint 

(a) 

(b) 

Figure III.4.1-15: Stress distribution for the axial test: (a) 18 mm thickness joint, and (b) 36 mm thickness joint 

(a) 

(b) 

(c) 

Figure III.4.1-16: Critical elements for the bending test: (a) pipe, (b) flange, and (c) bolt 

111 
 
 
 
 
 
 
 
 
 
(a) 

(b) 

(c) 

Figure III.4.1-17: Critical elements for the axial test: (a) pipe, (b) flange, and (c) bolt 

III.4.1.3.

Parametric numerical study of Tee pipe joints 

Figure III.4.1-18: Tee-junction capacity in (a) out-of-plane bending and (b) in plane bending. 

Figure  III.4.1-19:  Tee-junction  capacity  in  axial  loading  (a)  pulling  branch  away  from  the  main  pipe  and  (b) 
pushing branch towards the pipe. 

Figure III.4.1-20: Variation of yield and ultimate strength in for the 3 branch-to-pipe diameters in bending. 

020406080100120140020406080100Imposed displacement (mm)Force (kN)ReinforcedFitting(a)020406080100120140020406080100Imposed displacement (mm)Force (kN)ReinforcedFitting(b)0200400600800100012001400020406080100Imposed displacement (mm)Force (kN)ReinforcedFitting(a)02004006008001000020406080100Imposed displacement (mm)Force (kN)ReinforcedFitting(b)0501001502002503008x48x68x8pipe-to-branch diameterStrength (kN)Yield; out-of-plane bendingUltimum; out-of-plane bendingYield; in-plane bendingUltimum ; in-plane bending112 
 
 
 
 
 
 
 
 
 
 
Figure III.4.1-21: Comparison of out-of-plane and in-plane strength for the different pad geometries. 

III.5.  Appendix of WP5 

III.5.1.  SEISMIC RESPONSE OF LIQUID STORAGE TANKS 

Representative  results  of  the  behavior  of  Tank  I  are  displayed  in  the  present  Appendix.  Numerical 
results for the unanchored tank are examined first, and compared with results from the anchored tank. 
General-purpose finite element program ABAQUS is used for the simulation of the structural behavior 
of the tank under static loading, assuming anchored and unanchored conditions. 

For the case of unanchored conditions, the tank shell and the bottom plate are modeled with shell finite 
elements, with special attention on the annular base plate of the tank. More specifically, the tank and the 
bottom plate are modeled with four-node reduced-integration shell elements (S4R). In such a case, the 
effects of uplifting of the unanchored tank are simulated with the use of appropriate contact conditions 
between the bottom plate of the tank and the ground. The ground is simulated using solid eight-node 
reduced-integration  elements  (C3D8R).  The  friction  coefficient  between  the  tank  base  plate  and  the 
ground  is  assumed  to  be  0.3.  The  tank  roof  is  modeled  with  the  use  of  two-node  (linear)  beam-type 
finite elements (B31). 

An incremental non-linear static analysis is conducted in three steps. In the first step, gravity of the tank 
is applied, whereas in the second step, hydrostatic pressure of the liquid is activated. Finally, in the third 
step,  horizontal  hydrodynamic  loading  is  applied,  through  an  incremental  elastic-plastic  “pushover” 
analysis with large displacements. The distribution of hydrodynamic pressures on the tank wall follows 
the  corresponding  solution  of  the  hydrodynamic  problem,  for  both  the  impulsive  and  convective 
component  of  liquid  motion  It  is  applied  on  the  inner  surface  of  the  tank  shell  and  the  bottom  plate 
through a special-purpose load-user subroutine in the ABAQUS model. 

The  results  of  this  analysis  indicate  a  significant  effect  of  uplifting  in  tank  behavior,  when  compared 
with  the  case  of  anchored  tanks.  Those  results  are  aimed  at  better  understanding  the  uplifting 
phenomenon  during  seismic  excitation,  and  may  be  used  for  improvement  of  the  current  design 
practice. 

A non-linear dynamic analysis is also performed by simulating the tank-liquid system with a simplified 
and the 
model that accounts for the geometry of the liquid storage tank of radius 
hydrodynamic  response  of  the liquid. The  tank  is  assumed  to  be  excited  by  a  horizontal  base-ground 
seismic  acceleration  motion, 
.  In  the  case  of  anchored  tanks,  the  container  is  anchored  to  the 
foundation and the impulsive and convective motion of the liquid are modelled by assuming two linear 
oscillators with masses 
 respectively, as seen in Figure III.5.1-6(a). In 
and 
the case of unanchored tanks, the container lies on the foundation and its base plate without anchors. 
The main difference between anchored and unanchored tanks is the uplifting mechanism which can be 
modelled  by  adding  a  nonlinear  spring  at  the  tank  (Figure  III.5.1-6(b)).  For  accurate  modeling,  the 
relationship of Figure III.5.1-4 is used. Finally, to facilitate the formulation, the convective mass 

 and filling height 

 at heights 

and 

has not been considered. 

RHXImCmIhChM113 
 
 
(a) 

(b) 

Figure III.5.1-1: Finite element analysis of unanchored liquid storage tank under lateral loading; (a) uplifting of 
base plate, (b) development of elephant’s foot at the tank bottom. 

Figure  III.5.1-2:    Top  horizontal  displacement 
moment

, subjected to finite element pushover analysis. 

  and  maximum  vertical  uplift  of  Tank  I  versus  overturning 

Figure  III.5.1-3:    Ratio  of  maximum  compressive  axial  membrane  forces  for  the  unanchored over  the  anchored 
tank  versus  the  overturning  moment;  finite  element  results  for  Tank  I  and  curves  for 
0.8,  1.0,  2.0, 
adopted by EN 1998-4. 

uM/HR114 
 
 
 
 
 
 
 
 
 
 
Figure III.5.1-4:  Overturning moment in terms of the rotation of the tank. 

Figure  III.5.1-5:  Length of  uplifted part of  the base as a  function of the  vertical  uplift  at the edge and bilinear 
approximation to be used later in the dynamic analysis. The equations of the two lines are shown on the graph. 

(a) 

(b) 

Figure III.5.1-6: Simplified models for anchored and unanchored liquid storage tank. 

115 
 
 
 
 
 
 
Figure III.5.1-7: Time history of the uplifting axial membrane force in the case of (a) the anchored Tank I and (b) 
the unanchored Tank I for the earthquake in Duzce, Turkey (1999). 

The analysis of elephant’s foot buckling at the bottom of Tank I is examined, through the consideration 
of  an  axisymmetric  finite  element  model.  It  can  be  concluded  that  there  is  a  quasi-uniform  state  of 
deformation along a significant part of the tank perimeter, justifying the axisymmetry assumption. The 
most  important  observation  refers  to  post-buckling  response  of  the  tank,  which  appears  to  be  very 
unstable;  after  reaching  a  maximum  stress  (load)  the  capacity  drops  very  rapidly,  indicating  that  the 
tank has not the ability of absorbing significant amounts of inelastic energy. 

Figure III.5.1-8:  Development of elephant’s foot, as predicted by the finite element analysis. 

III.6.  Appendix of WP6 

III.6.1.  Task 6.1 Seismic behaviour of vertical-cylindrical pressure vessels 

Three different geometries are considered for the case of vertical-cylindrical slender pressure vessels on 
skirt  supports,  a  short,  a  tall  and  a  thick  vessel  as  described  in  Deliverable  6.1,  extending  the  work 
carried out in Deliverable 2 for case study 5. 

III.6.1.1.

Pushover analysis results discussion 

A non-linear static (pushover) analysis has been employed to identify the modes of failure and study the 
progressive collapse after exceeding the yield limit. For the case of the vertical-cylindrical vessel No1 
the  base  shear  versus  drift  is  plotted  in  Figure  III.6.1-1  for  both  the  rectangular  and  triangular 
assumption  for  the  seismic  load  distribution  as  well  as  the  effect  of  internal  pressure.  Maximum 
operating pressure (PS), design pressure (Pd) and zero internal pressure was considered. 

(a) (b) (c) elephant’s foot bucklingTank shellBottom plate00.050.10.150.20.250.30.3500.010.020.030.040.05normalized horizontal displacement (δ/R)normalized heightδ/L=0.00052δ/L=0.00126δ/L=0.0063D=27.4 mH=16.514 mHliq=15.7 m116 
 
 
 
 
 
 
 
 
 
The maximum capacity is reached very quickly after the first-yield point, whereas the vessel resistance 
is significantly reduced beyond the maximum loading stage; this indicates that the vessel may not be 
capable  of  absorbing  significant  amount  of  inelastic  energy.  The  effect  of  internal  pressure  is 
insignificant  up  to  the  point  of  maximum  capacity  of  the  vessel  but  is  beneficial  for  post-buckling 
behavior. By comparing the seismic forces considered in the initial design of these pressure vessels it 
was observed that these lie within the elastic behaviour of the pushover curve and there is still reserve 
to endure higher seismic loads. 

Figure III.6.1-1: Base shear versus roof displacement – pushover analysis results for the VV No1 

The corresponding mode of failure is buckling at the first shell course immediately above the skirt for 
vessels 1 and 2, whereas for the thick pressure vessel 3 the critical mode of failure is buckling on the 
skirt, as shown in Figure III.6.1-2. 

The  base  shear  forces  derived  from  the  pushover  analysis  are  plotted  in  Figure  III.6.1-3  for  vertical 
vessel  No3  together  with  the  results  of  IDA  analysis.  Pushover  analysis  using  the  triangular  load 
distribution  is  conservative,  while  the  rectangular  distribution  could  overestimate  the  capacity  of  the 
vessels. 

No1

No3

Figure III.6.1-2: Equivalent plastic strain at maximum capacity for VVs No1 and 2 

0100020003000400050000200400600800100012001400160018002000Drift (mm)Base shear (kN)P=0 triangular distributionP=PS triangular distributionP=Pd triangular distributionP=0 unifrorm distributionP=PS unifrorm distributionP=Pd unifrorm distributionShear Force 1.04MN q=2PEEQ 4.7%117 
 
 
 
 
 
Figure III.6.1-3: Comparison of base shear vs drift between pushover and IDA results for the VV No3 

III.6.1.2.

Incremental dynamic analysis results discussion 

The  vertical-cylindrical  vessels  have  been  analyzed  under  incremental  dynamic  analysis  in  order  to 
estimate the behavior factor (Deliverable 6.1). Based on the pushover analysis results presented in the 
previous section, two regions of the vessel have been identified as critical, namely (a) the skirt and (b) 
the 1st shell course. In order to estimate the 
 factor, two seismic input levels have to be identified in 
each of the response curves; (a) the level where transition from linear to non-liner behavior occurs (λe) 
and (b) the level at which a specific failure criterion is satisfied (λu). Three failure criteria have been 
used: (a) Rupture of vessel wall when the equivalent plastic strain reaches a value equal to 0.4%, (b) 
Buckling  of  vessel  wall  when  the  compressive  strain  in  the  longitudinal  direction  reaches  a  critical 
value  defined  by  [38]  and  (c)  Instability  when  the  maximum  value  of  the  principal  structural  strain 
reaches a value equal to 5% according to EN13445-3 Annex B [24]. 

 factor is given below in Table III.6.1-1. 
Using the three failure criteria, the estimated value of the 
The  first  two  failure  criteria,  predict  behavior  factor  values  somewhat  lower  than  the  value  of  2 
suggested by ASCE 7 [5]. However, the failure criteria suggested from EN13445-3 gives much higher 
values. The more realistic results are believed to be given by the empirical formulas for buckling of the 
vessel  wall  giving  lower  behavior  factor  for  the  thin  walled  vertical  pressure  vessels  but  predicting 
better the behavior factor for the case of buckling of the skirt.  

Table III.6.1-1: Estimated q factor by IDA for the vertical pressure vessels 

VV No1 
FC 2 
1.5 
2.25 
1.5 

FC 3 
1.5 
4 
2.7 

FC 1 
1.5 
2.5 
1.7 

VV No2 
FC 2 
2.5 
4 
1.6 

FC 1 
2.5 
3.5 
1.4 

λe 
λu 
q 

FC 3 
2.5 
6 
2.4 

FC 1 
1.5 
2.5 
1.7 

VV No3 
FC 2 
1.5 
3.5 
2.3 

FC 3 
1.5 
4.0 
2.7 

III.6.2.  Task 6.2 Seismic behaviour of horizontal-cylindrical pressure vessels 

Four different geometries are considered for the case of horizontal-cylindrical pressure vessel on saddle 
supports.  The  first  pressure  vessel  represents  case  study  6  extensively  discussed  in  Deliverable  1.3b. 
The  other  vessels  are  variations  with  regards  to  reinforcing  rings,  length  and  thickness  as  given  in 
Deliverable 6.1. In HV2 the reinforcing rings have been omitted, while in HV3 the length between the 
saddle supports has been doubled. For HV4 an increase in thickness as well as increase in weight of the 
vessel has been considered. 

qq0100002000030000400000100200300400Drift (mm)Base shear (kN)uniform distributiontriangular distributionIDA Akz1IDA Akz2IDA Akz3IDA Akz4IDA Akz5IDA Akz6IDA Akz7No 3118 
 
 
 
 
 
 
 
 
III.6.2.1.

Pushover analysis results discussion 

Force displacement diagrams (pushover curves) are depicted in Figure III.6.2-1 for all model variations 
in the longitudinal and transverse directions. The pushover analysis shows that for the level of seismic 
loads  used  for the  design of  those  vessels, the  behavior is  elastic; in fact, the level  of seismic  design 
loads is well below the ultimate capacity level indicated by the pushover curves. A “ductile” behavior is 
observed,  with  the  vessel  capable  of  sustaining  significant  deformation  in  the  inelastic  range.  The 
capacity of the vessel in the transverse direction is double that in the longitudinal. 

Figure III.6.2-1: Base shear versus roof displacement for all horizontal vessels with internal pressure  – pushover 
analysis results (a) in the longitudinal and (b) in the transverse direction. 

The  location  of  failure  is  given  in  Figure  III.6.2-2  for  one  representative  case  for  loading  in  the 
longitudinal and transverse directions respectively. It can be seen that the mode of failure is buckling at 
the saddle support. The exact location in the saddle varies with load direction as depicted in the figures. 

The  effect  of  internal  pressure  when  pushover  loading  is  applied  in  the  longitudinal  direction  is  not 
significant.  This  is  true  for  all  pressure  vessels  modeled.  Whereas,  when  loading  is  applied  in  the 
transverse  direction  the  effect  is  more  pronounced  for  no  internal  pressure.  This  can  be  explained 
because for loading in the longitudinal direction stress concentration is limited in the saddle, however 
when loading is applied in the longitudinal direction and there is no internal pressure then there is also 
stress concentration in the shell next to the saddle, making the effect of internal pressure significant. 

The reinforcing rings as well as increased thickness of the vessel increases the capacity of the vessel to 
endure seismic loading as expected. Reinforcing rings are recommended for large pressure vessels or 
pressure vessels located in high seismicity areas. Caution shall be taken when deciding the area of the 
saddle reinforcement as the adjacent shell area is expected to fail in the case of no internal pressure. A 
simplified  model  of  a  single  support  and  part  of  the  vessel  could  be  analysed  instead  of  the  full 
structure. 

(a)

(b)

Figure  III.6.2-2:  PEEQ  distribution  for  the  Horizontal  Vessel  No1  with  internal  pressure  –  (a)  loading  in  the 
longitudinal and (b) in the transverse direction 

020040060080010000102030405060708090100Drift (mm)Shear Force  (kN)HV1 PSHV2 PSHV3 PSHV4 PS(a)050010001500200025000102030405060708090100Drift (mm)Shear Force (kN)HV1 PSHV2 PSHV3 PSHV4 PS(b)119 
 
 
 
 
 
Figure  III.6.2-3: Effect of internal pressure  on base shear  versus roof displacement for HV2– pushover analysis 
results (a) in the longitudinal and (b) in the transverse direction 

III.6.3.  Task 6.3 Seismic behaviour of pressure vessels of spherical shape 

Two different types of pressure vessels (braced and non-braced pressure vessels with 12 Columns) have 
been  investigated  using  modal  and  pushover  analyses  as  well  as  incremental  dynamic  analyses. 
Furthermore more on the basis of both examples the investigations have been extended by a parametric 
numerical study with 78 different geometries of braced and non-braced 4-, 8- and 12-column systems. 

For the calculation of fundamental periods a simplified beam method were derived. Using this method 
an excellent agreement with the results of the performed modal analyses has been observed. 

III.6.3.1.

Pushover analysis results discussion 

For the pushover analyses the same FE-models were used as for the modal analyses with exception of 
some  crucial  details;  geometrical  non-linearities  and  inelastic  material  behaviour  in  terms  of  a  von 
Mises plasticity material model with linear isotropic hardening (slope of about 500 times smaller than in 
the elastic range), based on the nominal values of P355 GH were considered. In the first load step the 
structure  was  loaded  by  the  operating  load  including  the  dead  weight  and  the  maximum  weight  of 
filling  (kf  =  90  %);  and  in  the  second  step  the  structure  was  loaded  displacement-controlled  using  a 
sliding  boundary  condition  at  the  centre  of  gravity,  which  was  coupled  to  the  shell.  Furthermore  the 
influence of internal pressure on the structural behaviour of the vessel has been investigated. 

The results of the pushover analyses for the braced and non-braced system are given in Figure III.6.3-1 
in terms of a force-displacement curve (pushover curve) referred to the centre of gravity. 

Figure  III.6.3-1: Pushover analysis results in terms of force-displacement curve referred to the centre of gravity 
for the non-braced system (left) and the braced system (right). 

The  non-braced  system  behaved  elastic  without  any  significant  plastifications  up  to  a  displacement 
equal to 132 mm. The first plastifications occurred directly below the shell-column connections of the 
columns  which  are  loaded  under  the  greatest  compressive  stress  due  to  the  overturning  moment. 

020040060080010000102030405060708090100Drift (mm)Shear Force  (kN)HV2 PSHV2 P=0HV2 Pd(a)05001000150020000102030405060708090100Drift (mm)Shear Force (kN)HV2 PSHV2 P=0HV2 Pd(b)0.01.12.23.34.45.56.67.78.80120240360480600Base shear force [MN]Drift [mm]Operating load (kf = 90 %)Operating load + internal pressureYieldloadUltimateload0.03.26.49.612.816.019.222.425.60120240360480600Base shear force [MN]Drift [mm]Operating load (without imperfections and εpl,limit)Operating load (pre-curved braces)Operating load (pre-curved braces) + internal pressureYieldloadUltimateloadNon-braced PV120 
 
 
 
 
 
 
 
Thereafter  up  to  a  displacement  equal  to  265  mm  (ultimate  state)  all  columns  successively  plastified 
also  directly  below  the  shell-column  connections.  By  reaching  the  ultimate  state  first  local  buckling 
occurred  (see  Figure  III.6.3-2,  left).  And  finally  up  to  a  displacement  of  approx.  600  mm  all  other 
columns buckled locally, whereby for the columns loaded in tension due the overturning moment the 
location  of local  buckling  shifted in  the  range  of  compressive  strains  due to  bending  in  the  columns. 
Furthermore it should be mentioned that no plastifications occurred in the sphere up to the end of the 
complete loading history (1000 mm). 

260 mm displacement 
Non-braced PV 

134 mm displacement 
Braced PV 

Figure  III.6.3-2:  Pushover  analysis  results  in  terms  of  deformed  shape  of  critical  details  showing  failure  modes 
and equivalent plastic strains PEEQ. 

The braced system behaved elastic without any significant plastifications in columns or bracings up to a 
displacement equal to 45 mm. At the time of reaching the yield load the first plastifications occurred in 
the most loaded bracings, which were directed parallel to the load direction and therefore loaded under 
the greatest tension stress. Immediately after reaching the yield point all bracings in compression failed 
rather  quickly  by  formation  of  plastic  hinges.  Thereafter  up  to  a  displacement  equal  to  197  mm 
(ultimate state) the strongest loaded bracings in tension failed due to reaching of the ultimate elongation 
(see  Figure  III.6.3-2,  right),  which  leaded  in  consequence  to  the  first  plastifications  in  the  columns 
loaded  in  compression  due  the  overturning  moment.  After  reaching  the  ultimate  state  up  to  a 
displacement  equal  to  264  mm  all  bracings  in  tension  failed  as  well  as  all  columns  successively 
plastified directly below the shell-column connections, so that the post critical behaviour was equal to a 
non-braced system. 

III.6.3.2.

Incremental dynamic analysis results discussion 

In  order  to  examine  the  range  of  possible  behaviour factors  q,  which  could  be applied  to  braced  and 
non-braced  spherical  pressure  vessel  supporting  structures,  nonlinear  incremental  dynamic  analyses 
were  conducted  on  simplified  models.  In  these  simplified  models  the  pressure  vessel  sphere  was 
modeled  as  rigid  framework  since  pushover  analysis with the  detailed  model showed  no influence  of 
sphere  flexibility  up  to  the  displacements  of  interest.  A  set  of  10  artificial  accelerograms  using 
SIMQKE  program  is  considered  for  the  seismic  input.  The  selection  and  generation  of  the  artificial 
accelerograms takes into account the following objectives and requirements: 

1. comparability of time step calculations with practical design calculations using seismic spectrum 

2. consideration of seismic action randomness 

3. computational efficiency  

Incremental dynamic analyses have been conducted and appropriate failure criteria have been defined 
and  evaluated  to  determine  behaviour  factors.  A  conservative  estimate  using  the  most  critical 
earthquake  from  the  set  of  10  artificial  accelerograms  showed  that  the  behaviour  factors  q  are  larger 
than 2, what confirms current code provisions. 

121 
 
 
 
 
 
For  both  supporting  structures  dynamic  instability  was  evaluated  as  upper  limit  failure  criterion. 
Dynamic instability is defined as large increase of lateral drift of the structure with only a small increase 
in applied earthquake forces and was determined by means of Ballio-Setti method [7]. It was observed 
that in the case of non-braced systems severe dynamic instability could be reached suddenly whereby 
for the braced structure it occurred more smoothly and at higher earthquake intensities. The increase of 
lateral drift as function of seismic level for all ten accelerograms can be seen in Figure III.6.3-3 for both 
systems. 

(a) 

(b) 

Figure  III.6.3-3:  Maximum  lateral  displacements  versus  seismic  level  for  all  ten  accelerograms  for  (a)  the  non-
braced supporting structure and (b) the braced supporting structure 

Because the columns were modeled as fibre elements in the simplified model, failure criteria like local 
buckling  phenomena  could  not  directly  be  accounted  for.  Therefore  the  displacements,  which 
corresponded to these failure modes, were identified in the pushover analyses on the detailed model and 
then  used  as limit  displacements  in  the  incremental  dynamic  analyses.  For  the  non-braced  supporting 
structure the following failure criteria have been used: (a) dynamic instability according to Ballio-Setti 
method as mentioned above; (b)  achievement of 1st local buckling in column which corresponds well 
with achievement of ultimate load. This state occurred in the  pushover analysis with the detailed FE-
model at a displacement of 265 mm; (c) buckling of all columns. The accompanying displacement in 
the detailed model was determined to 363 mm. The resulting behaviour factors for each of the failure 
criterion are given in Table III.6.3-1 for the most critical earthquake as well as the calculated mean of all 
10 earthquakes. 

Table III.6.3-1: Most critical as well as mean behaviour factor for all failure criteria for the non-braced supporting 
structure 

Failure criterion 

Most critical q value 

Mean q value 

(a)  Dynamic instability 
(b)  1st buckling of any column 

(c)  Buckling of all columns 

3.0 

2.7 

3.3 

5.10 

3.30 

4.69 

Based on the observation that after the first column buckled, which was also the state of ultimate load in 
the detailed model, buckling of all other columns followed very quickly and was also followed quickly 
by severe and sudden dynamic instability problems in the simplified model, a behaviour factor q of 2.7 
is conservatively suggested for the non-braced supporting structure investigated herein. 

For  the  braced  supporting  structure  the  following  failure  criteria  have  been  used:  (a)  as  for  the  non-
braced  system  dynamic  instability;  (b)  acceptance  criterion  for  braces  in  tension  according  to  [35], 
where  the  maximum  plastic  elongation  of  bracings  is  limited  to  seven  times  the  axial  deformation  at 
tensile yielding load; (c)  full plastic behaviour of an arbitrary column head;  (d) limit displacement of 

00,20,40,60,811,21,4012345678910maximum lateral drift   umax[m]seismic levelλACC 1ACC 2ACC 3ACC 4ACC 5ACC 6ACC 7ACC 8ACC 9ACC 1000,20,40,60,811,21,4012345678910maximum lateral drift   umax[m]seismic level  λACC 1ACC 2ACC 3ACC 4ACC 5ACC 6ACC 7ACC 8ACC 9ACC 10122 
 
 
 
 
 
 
197 mm at which ultimate load was achieved in the pushover curve of the detailed model; and (e) first 
buckling of a column which was observed at an accompanying displacement of 284 mm The resulting 
behaviour factors for each criterion are listed in Table III.6.3-2. 

Table  III.6.3-2:  Most  critical  as  well  as  mean  behaviour  factor  for  all  failure  criteria  for  the  braced  supporting 
structure 

Failure criterion 

Most critical q value 

Mean q value 

(a)  Dynamic instability 

(b)  Plastic elongation of bracing 

(c)  Full column head plastification 

(d)  Ultimate load from Pushover 
(e)  1st buckling of any column 

3 to 4 

4.2 

2.3 

3.3 

3.9 

- 

5.77 

2.58 

3.94 

5.32 

Engineering judgment was necessary to define the dynamic instability failure criterion because although 
the maximum lateral drift was above the elastic reference line at almost all earthquake intensities, which 
would  result  in  failure  when  the  Ballio-Setti  method  was  strictly  applied,  no  sudden  increase  in 
deformation  occurred  up  to  seismic  levels  of  approximately  4  to  4.5.  Therefore  a  factor  of  4  seems 
reasonable for this failure criterion. If columns are designed and detailed in such a way, that sufficient 
ductility is ensured, preventing premature failure of columns due to strong plastification, a behaviour 
factor q of 3.3 is suggested for the investigated structure. Otherwise, if energy dissipation is desired to 
take place basically in the bracings, the q factor however should be restricted to a value of 2.3. 

123 
 
 
 
III.7.  Appendix of WP7 

III.7.1.  General analyses procedure for seismic design acc to EN 1998-1 and 4 

            Collection of design data:  - types of fluids  - pressures & temperatures  - piping data;  - seismic data - soil data - soil stratigraphy - deep geology (-determination seismic zones, IBC ground Maps)     Establish design seismic action criteria  -No-collapse requirement          (Ultimate Limit State):  PNCR, 50 Year  ≤ 10% or TNCR= 475 years* -Damage limitation requirement:  (serviceability limit state)  PDCR, 10 years≤ 10% or TDLR =  95 years *  *  Or: Target reliabilities, to be established by National Authorities   da           ag ≤ 0,04g (0,39 m/s2) or ag*S ≤ 0,05g (0,49 m/s2)         Design peak ground acceleration   ag =i* agR   ag =i* agR         Select method of  seismic analysis:              Equivalent static analyses            Modal reponse spectrum analyses     no yes           Evaluate consequences of failure:  - risk  to the public  -risk to the environment  - loss of investment - loss of production -  other ag ≤ 0,08g (0,78 m/s2) or ag*S ≤ 0,1g (0,98 m/s2)  yes no Determine ground type  Define Importance class (reliability differentiation by adjustment of return period TNCR Determine  Importance  factor I Reference elastic peak ground acceleration agR (type A ground,  TNCR= 475 years Very low seismicity area No seismic analyses acc. to EN 1998 required Low seismicity area simplified  seismic analyses possible Determine soil factor Soil type  → soil factor S* (response spectrum dependant!)  Time –history analyses Linear elastic analyses  Non-Linear  elastic-plastic analyses  124 
 
 
 
III.7.2. 

Seismic linear elastic analyses methods 

III.7.3. 

Investigation of seismic input 

III.7.3.1.

Shape of elastic response spectra 

Within the framework of the INDUSE project a comparison has been made between spectra derived for 
the  same  location  according  to  available  international  codes  (ASCE  7  vs  EN  1998-1  and  4.  Figure 
III.7.3-1 gives a comparison between spectra, generated for the same case, according to both codes: 

            Seismic analysis by Equivalent Static (uniform load) method : a)Response to uniform load:   (EN 13480-3) b)Response to Seismic Anchor Movement (SAM)      Modal elastic ground acceleration response spectrum  Representation of (horizontal and vertical) earthquake motion  at a given point at the surface) Same shape of response spectrum for both the ultimate- and the damage limitation state Seismic action described by  two independent  (horizontal) orthogonal components   da           Seismic dynamic analyses?         Design peak ground acceleration   ag =i* agR   ag =i* agR         Calculation of natural frequencies (possible modes of vibration) of the piping system:          Testing response against limit states   Calculation of system response   a) to modal elastic response spectra   b) to Seismic Anchor movement           Evaluation of system reponse  (SRSS method)     no             yes LRFD design  Evaluate type of structure - evaluation of non linear  energy dissipating capacity Determine  behaviour  factor  q(piping structure) q (substructure)  Seismic Dynamic analyses (with modal elastic response spectra (preliminary) Static analyses: a)Internal pressure b)Weight & sustained load  c)Flexibility   Amplification factor  for piping located at different levels  inside a structure   Design spectrum for elastic analyses  - Representation of (horizontal and vertical) earthquake motion  at a given point at the surface)  -  Building the piping model: Modeling end restraints (anchors, Tanks, Vessels) Mass distribution Support type and- spacing Component flexibility and stress intensification Number. of nodes between supports for  required accuracy of dynamic analyses   Testing response against limit states  Allowable stress design (ASD)  Ultimate limit states      da Seismic input Influence of sub structure  Modeling and analyses 125 
 
 
 
Figure III.7.3-1: Horizontal design spectrum data – EN 1998-1 vs ASCE 7 

Analyzing  Figure  III.7.3-1  it  can  be  observed  that  the  horizontal  design  spectra  generated  according 
ASCE  7  show  approximately  the  same  peak  level,  the  peak  of  the  ASCE  spectra  being  reached  at  a 
somewhat higher value of the response period and extending over a wider range of natural periods. One 
could  say  that  in  this  example  the  ASCE  7  spectra  cover  both  the  Type  1  and  type  2  spectra  of  the 
European code, except for very small periods.  

III.7.3.2.

Influence of natural frequency of the piping system 

The design acceleration to be considered in the calculation depends on both the shape of the response 
spectrum  and  the  stiffness  characteristics  (possible  modes  of  vibration)  of  the  piping  system  under 
consideration. This is illustrated in the calculation example of Annex A of the Guidelines: 

Figure III.7.3-2: Configuration of 10” ammonia transmission line. 

126 
 
 
 
 
 
 
Figure  III.7.3-2  shows  is  section  of  an  above  ground  10”,  25  bar  Ammonia  transportation  pipeline 
subject  to  considerable  temperature  difference,  supported  on  sliding  supports,  with  locally  side 
restraints. 

Figure  III.7.3-3: Horizontal and vertical seismic design spectra according EN 1998-1 and ASCE-7,  with natural 
frequency range (blue shaded) of above ground cross country ammonia pipeline system 

In  Figure  III.7.3-3,  the  blue  shaded  region  of  the  curve  represents  the  range  in  which  this  particular 
system’s  natural  frequencies  lie;  the  bold  lines  are  sample  modes  which  have  been  highlighted  (it  is 
observed  that  the  pipe  system  under  consideration  is  rather  flexible  ,  showing  relatively  low  natural 
frequencies). The region outside of the shaded area is above a certain cut-off period, which translates 
for  the  ammonia  line  to  0.866  Seconds,  which  is  1.154  Hz  –  the  1st  mode  of  the  systems  natural 
frequency. Although the peak of the type II horizontal curve is higher, the maximum acceleration with 
which  the  software  calculates  the  system  at  the  1st  mode  of  vibration  is  0.15  ms-2  for  the  type  I 
horizontal curve. The pipe’s natural frequency is the main factor on how much the spectra influence the 
results. 

III.7.4. 

Influence of supporting structure 

A key point in modeling a piping system is the possibility to neglect the interaction (static and dynamic) 
between the pipes and the supporting structure. A crude rule is based on the ratio WR of the weight of 
pipes  over  the  weight  of  the  supporting  structure.  In  particular,  if  WR  <  25%  the  interaction  can  be 
excluded  and  the  piping  system  can  be  considered  as  a  non-building  structure,  loaded  by  a  seismic 
action coming from the supporting structure at pipes level. This rule was investigated and found to be 
approximately useful. A recommendation has been taken up in the guidelines. 

 127 
 
 
 
 
 
Figure III.7.4-1: Piping system on a pipe-rack 

Figure III.7.4-2: Possible positions of pipes 

Figure III.7.4-3: Variation of the moment in the pipes for 
the coupling effect 

Figure III.7.4-4: Deformation of substructure 

III.7.5.  Example of ultimate limit state for elbows 

From the test series carried out at TU Delft and at CSM, a fatigue model was developed, showing very 
good conformance with the test results and able to predict the result of testing. From this a limit state 
for bends under strong cyclic loading can be defined. 

128 
 
 
 
 
 
 
 
 
 
 
Test No. Displacement range 

Number of cycles 

(mm) 

until failure 

Local strain 
range* (%) 

(Exp.) 

1 

2 

3 

4 

5 

6 

7 

8 

±25 

±70 

±100 

±150 

±200 

±250 

±300 

ECCS protocol 

13160 

444 

171 

61 

28 

17 

10 

16 

0.33 

1.23 

- 

2.61 

- 

3.84 

4.02 

- 

*: calculated in the elbow hoop direction at the critical location 

Table III.7.5-1: Displacement range and number of cycles 

Figure III.7.5-1: Increasing amplitude – Loading 
pattern according to ECCS recommendations. 

Figure III.7.5-2: Limit state for strong cyclic loading, as taken up in the design guideline 

lfN  129 
 
 
 
 
 
 
 
 
IV.  List of acronyms and abbreviations 

AF 

ASME   

CSM 

EC8 

ECCS   

FE 

LVDT   

MPS 

PGA 

RWTH  

SWOT   

TL 

Armstrong – Frederick model 

American Society of Mechanical Engineers 

Centro Sviluppo Materiall 

Eurocode 8 

European Convention for Constructional Steelwork  

Finite Element 

Linear Variable Differential Transformer 

Modal percentage of excited mass 

Peak Ground Acceleration 

Aachen University of Technology 

Strengths, Weaknesses, Opportunities, Threats 

Tseng-Lee 

TPLH / Technip 

TECHNIPETROL HELLAS SA 

TUDelft  

UThessaly  

UTrento 

WP  

Delft University of Technology 

University of Thessaly 

University of Trento 

Work Package 

elongation (%) 

diameter of tube 

elastic deformation of a specimen in the ECCS procedure for a cyclic test 

number of cycles to failure 

yield pressure 

tensile strength 

yield strength 

thickness  

displacement amplitude 

yield stress of a material 

“backstress” tensor  

material model parameter 

Young modulus  

equivalent plastic strain  

hardening modulus  

hardening modulus when two surfaces are in contact 

an appropriately calibrated parameter 

A%DyefNypmR0.%p2RtΔlyσaCEqHˆHh131 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
  
 
 
 
 
 
 
 
 
 
 
 
  
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
 
  
 
 
the intensification factor given in Appendix D of ASME B31.3 

initial yield stress 

mean size of the yield surface 

function that defines the size of the yield surface 

bending moment 

number of loading cycles  

number of cycles to failure 

deviatoric stress tensor 

allowable local stress amplitude  

basic allowable stresses at minimum metal temperature

basic allowable stresses at maximum metal temperature  

longitudinal stress  

longitudinal stress produced by occasional loads 

 pipe cross-section section modulus 

material model parameter 

initial distance between the two surfaces 

hoop strain range 

displacement amplitude 

yield stress of a material 

ii0kk()qkMNfNsASCShSLSL,OccSZinΔlyσ132 
 
 
 
 
 
 
 
 
 
 
   
 
  
 
 
 
 
 
 
 
 
 
 
 
 
 
 
  
 
 
  
  
 
 
 
 
 
   
 
 
 
 
 
 
 
  
 
 
 
 
 
 
 
 
 
 
 
 
V. List of figures 

Figure II.3.2-1: Test setup in the laboratory with test specimen P3-1. .................................................... 28 
Figure II.3.2-2: Schematic representation of the loading history according to ECCS [20] ..................... 28 
Figure II.3.3-1: Weld neck flange geometry ........................................................................................... 29 
Figure II.3.3-2: Test configuration, four point bending .......................................................................... 29 
Figure II.3.3-3: Testing device ................................................................................................................ 29 
Figure II.3.4-1: Experimental set-up: Front view .................................................................................... 32 
Figure II.3.4-2: Experimental set-up: Instrumentation ............................................................................ 32 
Figure II.3.4-3: Cracked elbow specimen configuration ......................................................................... 32 
Figure II.3.4-4: First load-displacement cycles for the 5 tests ................................................................ 33 
Figure II.3.4-5: Test 8 – ECCS load - displacement cycles .................................................................... 33 
Figure II.3.4-6: Specimen geometry and test configuration .................................................................... 33 
Figure II.3.4-7: Specimen inside the experimental device ...................................................................... 34 
Figure II.3.4-8: Specimen failure for cracking at flank position ............................................................. 34 
Figure II.3.4-9: Typical Force vs. displacement result, Test 2 ................................................................ 34 
Figure  II.3.5-1:  The  two  types  of  T-joint  branch  connection  studied:  (a)  fitting,  (b)  welded 
reinforcement ring-plate. ......................................................................................................................... 36 
Figure II.3.5-2:  Test setup for the T-joint tests: (a) overall view, (b) side view. ................................... 36 
Figure II.3.5-3: Deformed T-joint specimens at failure: (a) TJF, (b) TJWR........................................... 37 
Figure II.4.1-1: Material test results – elbow hoop direction .................................................................. 39 
Figure II.4.1-2: Lemaitre-Chaboche calibration method ......................................................................... 39 
Figure II.4.1-3: Calibration results – Test E1B T-4 ................................................................................ 39 
Figure II.4.1-4: Calibration results – Test E1B T-9 ................................................................................ 39 
Figure II.4.1-5: Calibration results – Test E1B T-5 ................................................................................ 39 
Figure II.4.1-6: Test on elbows – Simulation results using Set 1 and Set 3 ............................................ 39 
Figure II.4.1-7: A schematic view of the two-surface model. ................................................................. 41 
Figure II.4.1-8: Material curve and model predictions ............................................................................ 41 
Figure II.4.2-1: General view of the numerical model ............................................................................ 42 
Figure II.4.2-2: Concentrations of plastic deformations – closing bending loads ................................... 42 
Figure II.4.2-3:  Numerical simulation of Test 4 - Load-displacement cycles ........................................ 43 
Figure II.4.2-4:  Numerical simulation of Test 4 -  Flattening- displacement cycles .............................. 43 
Figure II.4.2-5:  Numerical simulation of Test 4 - Evolution of maximum local strains ........................ 43 
Figure II.4.2-6: Numerical simulation of Test 4 - Evolution of minimum local strains .......................... 43 
Figure II.4.2-7: Numerical simulation of Test No1-  Load - displacement ............................................. 44 
Figure II.4.2-8: Numerical simulation of Test No1 - Ovalization - displacement curve ......................... 44 
Figure II.4.2-9: Numerical simulation of Test No2: Load - displacement .............................................. 44 
Figure II.4.2-10: Concentration of strains at the elbow critical region ................................................... 44 
Figure II.4.2-11: Local hoop strain range versus number of cycles to failure......................................... 45 
Figure II.4.2-12: The piping system and the two substructures .............................................................. 46 
Figure II.4.2-13: Hardware-Software configuration for the PDDS ......................................................... 47 
Figure II.4.2-14: Displacement histories and relevant spectra of the Coupling Node #1 at (a) SLVT and 
(b) SLCT (Blue: Experimental; Magenta: Numerical) ............................................................................ 47 
Figure II.4.2-15: FE-models overview: (P1) internal-, (P2) external- and (P3) nozzle-reinforcement. .. 48 
Figure II.4.2-16: Mesh details ................................................................................................................. 49 
Figure II.4.2-17: Monotonic and cyclic material properties employed in FEA ...................................... 50 
Figure II.4.2-18: PN63-1 FE model deformed shape and plastic strains at maximum bending deflection 
(right) and specimen at the end of the test (left) ...................................................................................... 50 
Figure II.4.2-19: PN63-1 FE model deformed shape and plastic strains at maximum bending deflection 
(right) and specimen at the end of the test (left) ...................................................................................... 50 
Figure II.4.2-20: Geometrical models for the: (a) bending test, and (b) axial test .................................. 51 
Figure  II.4.2-21:  Comparison  between  experimental  and  numerical  data  for  bending  tests:  (a)  Design 
01, (b) Design 02; and axial tests: (c) Design 01..................................................................................... 52 
Figure II.4.2-22: Deformed shapes and stress levels in the joint for the: (a) bending test – Design 01 –, 
(b) bending test – Design 02 –, and (c) axial test – Design 01 –. ............................................................ 52 
Figure II.4.2-23: Numerical models: General view ................................................................................. 53 

133 
Figure II.4.2-24: Finite element analysis results:  TJF 2C test ................................................................ 54 
Figure II.4.2-25: Numerical TJF model .................................................................................................. 54 
Figure II.4.2-26: Finite element analysis results:  TJWR 2C test ............................................................ 54 
Figure II.4.2-27: Numerical TJWR model .............................................................................................. 54 
Figure II.4.3-1: Numerical model for the elbows used for the parametric analysis ................................ 56 
Figure II.4.3-2:Numerical parametric analysis results – 8-inch SCH40 long radius - Monotonic analysis
 ................................................................................................................................................................. 56 
Figure II.4.3-3: General view of the numerical model ............................................................................ 56 
Figure II.4.3-4: Concentrations of plastic deformations ......................................................................... 56 
Figure II.4.3-5:Moment versus end-rotation curve for pressure 0, 3.2 and 7.0 MPa .............................. 57 
Figure II.4.3-6: Moment versus end-rotation curve for displacement ranges ±0.05, ±0.1, ±0.25 and ±0.4 
rad, (pressure 0 MPa). ............................................................................................................................. 57 
Figure II.4.3-7: Overview of the geometrical parameters for the parametric study ................................ 57 
Figure II.4.3-8: FE model and load scheme considered in parametric study .......................................... 60 
Figure II.4.3-9: Comparison among different thickness flanges for bending tests ................................. 61 
Figure II.4.3-10: Comparison among different thickness flanges for axial tests ..................................... 61 
Figure  II.4.3-11:  Stress  distribution  for  the  bending  test:  (a)  18  mm  thickness  joint,  and  (b)  36  mm 
thickness joint .......................................................................................................................................... 62 
Figure  II.4.3-12:  Stress  distribution  for  the  axial  test:  (a)  18  mm  thickness  joint,  and  (b)  36  mm 
thickness joint .......................................................................................................................................... 63 
Figure II.4.3-13: Critical elements for the bending test: (a) pipe, (b) flange, and (c) bolt ...................... 63 
Figure II.4.3-14: Critical elements for the axial test: (a) pipe, (b) flange, and (c) bolt ........................... 63 
Figure II.4.3-15: Reinforcement pad (EN13480-3 [25]) and tee-fitting geometry (EN10253-2 [22]). ... 64 
Figure II.5.2-1: The three liquid storage tanks used in the design examples. ................................... 69 
Figure II.7.1-1: The piping system considered in the present study. ....................................................... 71 
Figure III.1.2-1: Three liquid storage tanks which are used in the design examples .............................. 78 
Figure III.1.3-1: Sketch of (a) vertical pressure vessel on skirt support and (b) horizontal pressure vessel 
on saddle supports (drawings made by software PV Elite 2010). ........................................................... 78 
Figure III.1.4-1: The analyzed piping system on a pipe-rack .................................................................. 79 
Figure III.1.4-2: Configuration of 10” ammonia transmission line. ........................................................ 79 
Figure III.1.4-3: Configuration of 20” gas transmission pipeline. .......................................................... 79 
Figure III.2.1-1:  Tank I (left), tank II (center) and tank III (right) used for the parametric studies in the 
present paper............................................................................................................................................ 79 
Figure  III.2.1-2:  Finite  element  analysis  of  unanchored  liquid  storage  tank  under  lateral  loading;  (a) 
uplifting of base plate, (b) development of elephant’s foot at the tank bottom. ...................................... 80 
Figure III.2.1-3: Elephant’s foot buckling, as obtained by finite element simulation. ............................ 80 
Figure  III.2.2-1:  Von  Mises  stress  distribution  at  maximum  capacity  (approximately  300mm 
displacement)........................................................................................................................................... 81 
Figure III.2.2-2: Von Mises stress distribution for the horizontal pressure vessel in the (a) longitudinal 
x-direction  (approx.  10  mm  displacement)  and  (b)  transverse  z-direction    (approx.  50  mm 
displacement)........................................................................................................................................... 81 
Figure III.3.1-1: API X52 from elbows cyclic properties at various imposed strain levels, longitudinal 
direction (left) transverse direction (right) .............................................................................................. 82 
Figure III.3.1-2: API X52 from elbows cyclic properties at various imposed stress levels, longitudinal 
direction (left) transverse direction (right) .............................................................................................. 83 
Figure III.3.1-3: P355 from elbows cyclic properties at various imposed stress levels .......................... 83 
Figure III.3.1-4: Stress amplitude versus number of cycles until failure. ............................................... 84 
Figure III.3.2-1: Load directions for monotonic loading......................................................................... 85 
Figure III.3.2-2: P1-1, P2-1 and P3-1– Cyclic load-displacement curve and pictures of the specimens at 
the end of the loading history (longitudinal load direction) .................................................................... 86 
Figure III.3.2-3: P1-2, P2-2 and P3-2– Cyclic load-displacement curve and pictures of the specimens at 
the end of the loading history (transverse load direction) ....................................................................... 87 
Figure III.3.3-1: Strain gauges position mapping .................................................................................... 88 
Figure III.3.3-2: Measured deflection at mid sections and rotation at load application points ............... 88 
Figure III.3.3-3: Actuator load vs. displacement Test PN63-1 ............................................................... 89 
Figure III.3.3-4: PN63-1 specimen after testing ...................................................................................... 89 

134 
Figure III.3.3-5: Actuator load vs. displacement Test PN40-1 ............................................................... 89 
Figure III.3.3-6: PN40-1 specimen after testing ...................................................................................... 89 
Figure III.3.3-7: Actuator load vs. displacement Test PN63-2 ............................................................... 89 
Figure III.3.3-8: PN63-2 specimen after testing ...................................................................................... 89 
Figure III.3.3-9: Actuator load vs. displacement Test PN63-3 ............................................................... 90 
Figure III.3.3-10: PN63-3 specimen after testing .................................................................................... 90 
Figure III.3.3-11: Actuator load vs. displacement Test PN40-2 ............................................................. 90 
Figure III.3.3-12: PN40-2 specimen after testing .................................................................................... 90 
Figure III.3.3-13: Actuator load vs. displacement Test PN40-3 ............................................................. 90 
Figure III.3.3-14: PN40-3 specimen after testing .................................................................................... 90 
Figure III.3.3-15: Actuator load vs. displacement Test PN40-4 ............................................................. 90 
Figure III.3.3-16: PN40-4specimen after testing ..................................................................................... 90 
Figure  III.3.3-17:  (a)  Designed  non-standard  BFJs  and  dimensions;  (b)  spiral  wound  gasket;  (c)  stud 
bolt ........................................................................................................................................................... 91 
Figure  III.3.3-18:  (a)  Bending  test  specimen,  BS1;  (b)  axial  test  specimen,  AS1;  (c)  a  spiral  wound 
gasket; (d) a stud bolt. ............................................................................................................................. 91 
Figure III.3.3-19: Test set-up for (a) bending tests; (b) axial tests. ......................................................... 92 
Figure  III.3.3-20:  (a)  Actuator  load-displacement  curve  from  test-  BSML18;  moment-rotation  curves 
from (b) test- BSCL18, (c) test- BSML27 and (d) test- BSCL27. .......................................................... 92 
Figure  III.3.3-21:  Load-flange  displacement  curve  from  (a)  test  ASML18,  (b)  test  ASCL18,  (c)  test 
ASCL27-1 and (d) test ASCL18-2. ......................................................................................................... 93 
Figure III.3.3-22: Failure of the pipe and bending of a bolt during test BSCL27 ................................... 93 
Figure III.3.4-1: Cracked elbow specimen No. 4 with displacement range ±150 mm: Failure stage with 
through thickness crack development ..................................................................................................... 94 
Figure III.3.4-2: Experimental set-up: Instrumentation ........................................................................... 94 
Figure III.3.4-3: Increasing amplitude – Loading pattern according to ECCS recommendations. ......... 95 
  Load – displacement cycles ...................................................... 95 
Figure III.3.4-4: Test 1 - 
 Load – displacement cycles ....................................................... 95 
Figure III.3.4-5: Test 2 - 
 Load – displacement cycles ..................................................... 95 
Figure III.3.4-6: Test 3 - 
Figure III.3.4-7: Test 4 - 
  Load – displacement cycles .................................................... 95 
 Flattening – displacement cycles ............................................. 95 
Figure III.3.4-8: Test 4 - 
 Load – displacement ............................................................... 96 
Figure III.3.4-9: Test 5 -  
 Load – displacement .............................................................. 96 
Figure III.3.4-10: Test 6 - 
 Load – displacement cycles ................................................... 96 
Figure III.3.4-11: Test 7 - 
Figure III.3.4-12: Test 7 - 
 Flattening – displacement cycles ........................................... 96 
Figure III.3.4-13: CSM FSTM in full scale testing machine ................................................................... 96 
Figure III.3.4-14: Measure of elbow displacements ............................................................................... 97 
Figure III.3.4-15: Measure of ovalization ............................................................................................... 97 
Figure III.3.4-16: Strain gauges installation ............................................................................................ 97 
Figure III.3.4-17: Test 1 load displacement curve (left) and final crack position (right) ........................ 97 
Figure III.3.4-18: Test 2 load displacement curve (left) and final crack position (right) ........................ 98 
Figure III.3.4-19: Test 3 load displacement curve (left) and final crack position (right) ........................ 98 
Figure III.3.4-20: Test 4 load displacement curve (left) and final crack position (right) ........................ 98 
Figure III.3.4-21: Test 5 load displacement curve (left) and final crack position (right) ........................ 99 
Figure III.3.4-22: Test 2 elbow ovalization ............................................................................................. 99 
Figure III.3.4-23: Test 4 elbow ovalization ............................................................................................. 99 
Figure III.3.4-24: (a) A 3D model of the piping system and support structure; (b) PS, NS and coupling 
nodes ...................................................................................................................................................... 101 
Figure III.3.4-25: Schematic of the experimental set-up for the hybrid tests ........................................ 101 
Figure III.3.4-26: Actual test set-up for the PDDS................................................................................ 101 
Figure III.3.4-27: Some components of the test specimen .................................................................... 102 
Figure III.3.4-28: Rotation histories and relevant spectra of the two end nodes of Elbow #1 at (a) SLVT 
and (b) SLCT ......................................................................................................................................... 102 
Figure III.3.4-29: Strain histories of Elbow #2 at (a) SLVT and (b) SLCT .......................................... 102 

25 lmm70 lmm100 lmm150 lmm150 lmm200 lmm250 lmm300 lmm300 lmm135 
Figure III.3.4-30: Displacement histories and relevant spectra of the Coupling Node #1 at (a) SLVT and 
(b) SLCT (Blue: Experimental; Magenta: Numerical) .......................................................................... 103 
Figure III.3.5-1: Load vs. LPD curves for the T-joint specimens under monotonic loading. ............... 104 
Figure III.3.5-2: Load vs. LPD curves under cyclic loading at an imposed LPD value of δ1=±60 mm: 
(a) TJF, (b) TJWR specimen. ................................................................................................................ 104 
Figure III.3.5-3: Load vs. LPD curves under cyclic loading at an imposed LPD value of δ2=±45 mm: 
(a) TJF, (b) TJWR specimen. ................................................................................................................ 105 
Figure  III.3.5-4:  Load  vs.  LPD  curves  under  cyclic  loading  according  to  the  ECCS  No45  testing 
procedure: (a) TJF, (b) TJWR specimen. .............................................................................................. 105 
Figure  III.3.5-5:  Through-thickness  crack  failure  of  TJWR  specimen  at  the  weld  toe  of:  (a)  branch 
region, (b) main pipe region at the end of the welded reinforcement plate. .......................................... 105 
Figure  III.3.5-6:  Through-thickness  crack  failure  in  the  region  of  the  fitting  for  the  TJF  specimen 
subjected to monotonic load. ................................................................................................................. 106 
 .. 106 
Figure III.4.1-1: Numerical parametric analysis results – 8-inch SCH40 long radius ,
 .. 106 
Figure III.4.1-2: Numerical parametric analysis results – 8-inch SCH40 long radius ,
 .. 106 
Figure III.4.1-3: Numerical parametric analysis results – 8-inch SCH40 long radius, 
 .. 106 
Figure III.4.1-4: Numerical parametric analysis results – 8-inch SCH40 long radius, 
Figure  III.4.1-5:  Moment versus  end-rotation  curve  for  displacement range ±0.25  rad,  (for pressure  0 
MPa, 3.2 MPa and 7.0 MPa). ................................................................................................................ 108 
Figure  III.4.1-6:  Moment  versus  end-rotation  curve  for  displacement  range  ±0.4  rad,  (for  pressure  0 
MPa, 3.2 MPa and 7.0 MPa). ................................................................................................................ 108 
Figure III.4.1-7: Moment versus end-rotation curve for displacement ranges ±0.05, ±0.1, ±0.25 and ±0.4 
rad, (pressure 0 MPa). ........................................................................................................................... 108 
Figure III.4.1-8: typical deformed shape and plastic strain map of the flanged joints after bending .... 109 
Figure III.4.1-9: Moment vs. rotation curves for monotonic bending of PN40 flanged connections ... 109 
Figure III.4.1-10: Moment vs. rotation curves for monotonic bending of PN63 flanged connections . 109 
Figure III.4.1-11: Moment vs. rotation curves for monotonic bending of PN100 flanged connections 110 
Figure III.4.1-12: Effect of pressure for all flange geometries .............................................................. 110 
Figure III.4.1-13: Effect of axial load for all flange geometries ........................................................... 110 
Figure  III.4.1-14:  Stress  distribution  for  the  bending  test:  (a)  18  mm  thickness  joint,  and  (b)  36  mm 
thickness joint ........................................................................................................................................ 111 
Figure  III.4.1-15:  Stress  distribution  for  the  axial  test:  (a)  18  mm  thickness  joint,  and  (b)  36  mm 
thickness joint ........................................................................................................................................ 111 
Figure III.4.1-16: Critical elements for the bending test: (a) pipe, (b) flange, and (c) bolt ................... 111 
Figure III.4.1-17: Critical elements for the axial test: (a) pipe, (b) flange, and (c) bolt ........................ 112 
Figure III.4.1-18: Tee-junction capacity in (a) out-of-plane bending and (b) in plane bending. .......... 112 
Figure III.4.1-19: Tee-junction capacity in axial loading (a) pulling branch away from the main pipe and 
(b) pushing branch towards the pipe. .................................................................................................... 112 
Figure  III.4.1-20:  Variation  of  yield  and  ultimate  strength  in  for  the  3  branch-to-pipe  diameters  in 
bending. ................................................................................................................................................. 112 
Figure III.4.1-21: Comparison of out-of-plane and in-plane strength for the different pad geometries.113 
Figure  III.5.1-1:  Finite  element  analysis  of  unanchored  liquid  storage  tank  under  lateral  loading;  (a) 
uplifting of base plate, (b) development of elephant’s foot at the tank bottom. .................................... 114 
Figure  III.5.1-2:    Top  horizontal  displacement 
  and  maximum  vertical  uplift  of  Tank  I  versus 
, subjected to finite element pushover analysis. .............................................. 114 
overturning moment
Figure III.5.1-3:  Ratio of maximum compressive axial membrane forces for the unanchored over the 
anchored  tank  versus  the  overturning  moment;  finite  element  results  for  Tank  I  and  curves  for 
0.8, 1.0, 2.0, adopted by EN 1998-4. .................................................................................. 114 
Figure III.5.1-4:  Overturning moment in terms of the rotation of the tank. ......................................... 115 
Figure III.5.1-5:  Length of uplifted part of the base as a function of the vertical uplift at the edge and 
bilinear  approximation  to  be  used  later  in  the  dynamic  analysis.  The  equations  of  the  two  lines  are 
shown on the graph. ............................................................................................................................... 115 
Figure III.5.1-6: Simplified models for anchored and unanchored liquid storage tank......................... 115 
Figure III.5.1-7: Time history of the uplifting axial membrane force in the case of (a) the anchored Tank 
I and (b) the unanchored Tank I for the earthquake in Duzce, Turkey (1999). ..................................... 116 

Δr=±0.02 radΔr=±0.04 radΔr=±0.08 radΔr=±0.12 raduM/HR136 
Figure III.5.1-8:  Development of elephant’s foot, as predicted by the finite element analysis. ........... 116 
Figure III.6.1-1: Base shear versus roof displacement – pushover analysis results for the VV No1 .... 117 
Figure III.6.1-2: Equivalent plastic strain at maximum capacity for VVs No1 and 2 ........................... 117 
Figure III.6.1-3: Comparison of base shear vs drift between pushover and IDA results for the VV No3
 ............................................................................................................................................................... 118 
Figure III.6.2-1: Base shear versus roof displacement for all horizontal vessels with internal pressure – 
pushover analysis results (a) in the longitudinal and (b) in the transverse direction. ............................ 119 
Figure III.6.2-2: PEEQ distribution for the Horizontal Vessel No1 with internal pressure – (a) loading in 
the longitudinal and (b) in the transverse direction ............................................................................... 119 
Figure III.6.2-3: Effect of internal pressure on base shear versus roof displacement for HV2– pushover 
analysis results (a) in the longitudinal and (b) in the transverse direction ............................................ 120 
Figure III.6.3-1: Pushover analysis results in terms of force-displacement curve referred to the centre of 
gravity for the non-braced system (left) and the braced system (right). ................................................ 120 
Figure III.6.3-2: Pushover analysis results in terms of deformed shape of critical details showing failure 
modes and equivalent plastic strains PEEQ. ......................................................................................... 121 
Figure III.6.3-3: Maximum lateral displacements versus seismic level for all ten accelerograms for (a) 
the non-braced supporting structure and (b) the braced supporting structure ....................................... 122 
Figure III.7.3-1: Horizontal design spectrum data – EN 1998-1 vs ASCE 7 ........................................ 126 
Figure III.7.3-2: Configuration of 10” ammonia transmission line. ...................................................... 126 
Figure III.7.3-3: Horizontal and vertical seismic design spectra according EN 1998-1 and ASCE-7, with 
natural frequency range (blue shaded) of above ground cross country ammonia pipeline system ....... 127 
Figure III.7.4-1: Piping system on a pipe-rack ...................................................................................... 128 
Figure III.7.4-2: Possible positions of pipes .......................................................................................... 128 
Figure III.7.4-3: Variation of the moment in the pipes for the coupling effect ..................................... 128 
Figure III.7.4-4: Deformation of substructure ....................................................................................... 128 
Figure III.7.5-1: Increasing amplitude – Loading pattern according to ECCS recommendations. ....... 129 
Figure III.7.5-2: Limit state for strong cyclic loading, as taken up in the design guideline .................. 129 

137 
 
 
 
VI.  List of tables 

Table II.3-1: Summary of experimental tests……………………………………………………….26 
Table II.3.1-1: Small scale testing performed to support full scale activity and FEA ............................ 27 
Table II.3.2-1: Number of cycles until failure Nf for longitudinal and transverse load direction ........... 28 
Table II.3.3-1: CSM test program on standard bolted flanged joints and main results / observations .... 30 
Table II.3.4-1: Displacement range and number of cycles ...................................................................... 32 
Table II.3.4-2: Specimen properties ........................................................................................................ 33 
Table II.3.4-3: Full scale test program .................................................................................................... 34 
Table II.3.4-4: Summary of Elbows test results performed by CSM ...................................................... 34 
Table II.3.5-1: Test results. ..................................................................................................................... 36 
Table II.4.1-1: Material model parameter values .................................................................................... 40 
Table II.4.2-1: Displacement range, number of cycles and local strain range. ....................................... 43 
Table II.4.2-2: Displacement range and number of cycles ...................................................................... 44 
Table II.4.2-3: Experimental and numerical results for Nf and Pf ........................................................... 49 
Table II.4.2-4: Error between experimental results and numerical simulations calculated with NRMSD 
and RMSD methods ................................................................................................................................ 52 
Table II.4.2-5: Experiments results ......................................................................................................... 53 
Table II.4.3-1: Geometrical characteristics of examined cases ............................................................... 55 
Table II.4.3-2: Geometrical parameters for the parametric study ........................................................... 57 
Table II.4.3-3: Parametric numerical results for cyclic loading in longitudinal load direction ............... 58 
Table II.4.3-4: Parametric numerical results for cyclic loading in transverse load direction .................. 58 
Table II.4.3-5: Flanged joints considered in parametric study ................................................................ 60 
Table II.4.3-6: Load cases analyzed for each geometry .......................................................................... 60 
Table II.4.3-7: Fracture load for bending tests with SMCS criterion ...................................................... 62 
Table II.4.3-8: Fracture load for bending tests with SMCS criterion ...................................................... 62 
Table III.3.1-1: Monotonic tensile properties of API X52 specimen extracted from elbow ................... 82 
Table III.3.1-2: Monotonic tensile properties of API5L X52 specimens extracted from pipes to be used 
in elbow piping specimens ...................................................................................................................... 82 
Table III.3.1-3: Summary of monotonic tensile properties of P355N extracted from straight pipes ...... 82 
Table III.3.1-4: Summary of monotonic tensile properties of P355 extracted from plates ..................... 83 
Table III.3.1-5: Summary of monotonic tensile properties of the base material. .................................... 84 
Table III.3.2-1: Yield loads and corresponding displacements of the internal-, external- and thickened 
nozzle reinforcement for different load directions .................................................................................. 85 
Table III.3.3-1: Dimensions of the designed non-standard BFJs (mm) .................................................. 91 
Table III.3.3-2:  Test program on non-standard BFJ carried out by UTrento ......................................... 91 
Table III.3.3-3: Comparison between experimental leakage moment and allowable moments suggested 
by codes ................................................................................................................................................... 94 
Table  III.3.3-4:  Comparison  between  experimental  leakage  load  and  allowable  loads  suggested  by 
codes ........................................................................................................................................................ 94 
Table  III.3.3-5:  Maximum  moment,  Mmax,  axial  force,  Nmax  and  shear  force,  τmax  in  the  piping 
system obtained from a case study .......................................................................................................... 94 
Table III.3.4-1:FSTM main capabilities .................................................................................................. 96 
Table III.3.4-2: Experimental program on a piping system carried out by UTrento ............................. 100 
Table III.3.4-3: PGA corresponding to different limit states ................................................................ 100 
Table  III.3.4-4:  Maximum  moment,  Mmax  stress,  σmax  and  shear,  τmax  in  the  piping  system  under 
reference seismic loading ...................................................................................................................... 100 
Table III.4.1-1: Parametric analysis results: 8-inch SCH20 long radius ............................................... 107 
Table III.4.1-2: Parametric analysis results: 8-inch SCH40 long radius ............................................... 107 
Table III.4.1-3: Parametric analysis results:  8-inch SCH80 long radius .............................................. 107 
Table III.4.1-4: Parametric analysis results: 8-inch SCH40 short radius .............................................. 107 
Table III.4.1-5: Parametric analysis results: 8-inch SCH20 long and short radius. .............................. 108 
Table III.4.1-6: Parametric analysis results: 8-inch SCH40 long and short radius. .............................. 108 
Table III.4.1-7: Parametric analysis results: 8-inch SCH20 long radius. .............................................. 108 
Table III.4.1-8: Summary of buckling moment and rotation for monotonic behavior. ......................... 110 
Table III.6.1-1: Estimated q factor by IDA for the vertical pressure vessels ........................................ 118 

139 
Table III.6.3-1: Most critical as well as mean behaviour factor for all failure criteria for the non-braced 
supporting structure ............................................................................................................................... 122 
Table  III.6.3-2:  Most  critical  as  well  as  mean  behaviour  factor  for  all  failure  criteria  for  the  braced 
supporting structure ............................................................................................................................... 123 
Table III.7.5-1: Displacement range and number of cycles .................................................................. 129 

140 
 
 
 
VII.  References 

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141 
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143 
 
European Commission

EUR 26319 —  Structural safety of industrial steel tanks, pressure vessels and piping systems under 

seismic loading (INDUSE)

Luxembourg: Publications Office of the European Union

2013 — 143 pp. — 21 × 29.7 cm

ISBN 978-92-79-34575-3
doi:10.2777/49423

145