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CRC Project: Smart Linings for Pipe and Infrastructure
State of The Art Literature Review on CIPP liners
Civil Engineering
AUTHORS: Guoyang Fu, Benjamin Shannon, Suranji Rathnayaka, Ravin Deo, Jayantha Kodikara Date: 02 / 04 / 2020
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QUALITY INFORMATION Document: Monash University CRC-P literature review CIPP
Edition date: 02-04-2020
Edition number: 1.7
Prepared by: Guoyang Fu
Reviewed by: Benjamin Shannon
Revision history
Revision Revision date Details Revised by
1 28-09-2018 Edited Benjamin Shannon
2 24/10/2018 Checked Suranji Rathnayaka
3 04/03/2019 Review Ravin Deo
4 07/03/2019 Final review and edit Benjamin Shannon
5 15/05/2019 Edit Guoyang Fu
6 23/09/2019 Revised Guoyang Fu
7 02/04/2020 Final Benjamin Shannon
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Contents
1. Introduction ........................................................................................................................ 1
1.1 Background ................................................................................................................. 1
1.2 History of CIPP liners ................................................................................................. 2
1.3 Current status in Australia ........................................................................................... 3
2. Liner type classification ..................................................................................................... 4
2.1 Class I .......................................................................................................................... 4
2.2 Class II......................................................................................................................... 4
2.3 Class III ....................................................................................................................... 5
2.4 Class IV ....................................................................................................................... 5
3. CIPP Liner types ................................................................................................................ 6
3.1 InsituMain®/Insituform .............................................................................................. 6
3.2 Aqua-Pipe®/ Sanexen .................................................................................................. 7
3.3 Starline®/Karl Weiss Technologies ............................................................................ 8
3.4 Saniline® W/ Sanivar ................................................................................................... 9
3.5 Installation of liners ..................................................................................................... 9
4. Liner imperfections, quality standards and installation risks ........................................... 10
5. Design methodology ........................................................................................................ 10
5.1 Current methodologies in standards .......................................................................... 10
5.2 Gap identification ...................................................................................................... 12
6. Review on Experimental Testing ..................................................................................... 12
6.1 Short-term tests ......................................................................................................... 12
6.1.1 Specimen tests .................................................................................................... 13
6.1.2 Large-scale pipe tests ......................................................................................... 31
6.2 Long-term tests .......................................................................................................... 38
6.2.1 Long-term pipe pressure testing ......................................................................... 38
6.2.2 Standard creep testing ........................................................................................ 41
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6.2.3 Accelerated creep testing ................................................................................... 42
6.2.4 Mechanical aging ............................................................................................... 46
6.3 Summary ................................................................................................................... 47
6.3.1 Short-term testing............................................................................................... 47
6.3.2 Long-term testing ............................................................................................... 56
6.4 Gap identification ...................................................................................................... 57
6.4.1 Short-term testing............................................................................................... 57
6.4.2 Long-term testing ............................................................................................... 59
7. Review of Numerical Modelling of corroded host pipes with imperfect liners .............. 59
7.1 Effect of size and geometry of a defect on host pipe ................................................ 59
7.2 Effect of material properties of the host pipe on the pressure rating of liners .......... 60
7.3 Effect of liner imperfections on the performance of the liners ................................. 61
7.4 Effect of ground movement ....................................................................................... 62
7.5 Effect of creep ........................................................................................................... 63
7.6 Summary ................................................................................................................... 64
7.7 Gap identification ...................................................................................................... 66
8. Conclusions, research gaps, and future research ............................................................. 66
8.1 Conclusions ............................................................................................................... 66
8.2 Research gaps ............................................................................................................ 67
8.3 Future research .......................................................................................................... 68
9. Acknowledgements .......................................................................................................... 70
10. References ..................................................................................................................... 70
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1. Introduction 1.1 Background
Deterioration of buried water and sewer pipes is a significant concern among many utilities in
Australia, North America, and other parts of the world. A number of different techniques have
been used to renew ageing pipelines that consist of a range of pipe materials. These techniques
include open trench replacements, replacement on new routes, pipe pulling and pipe bursting
(Morrison et al. 2013; Deb et al. 2015). The common type of pipe replacement method used in
Australia is open trench replacements and this method includes cutting and breaking of surface
materials and excavation of soil from the point of connection to the main along the entire length
of pipe to be replaced. However, this method of replacement is very expensive when replacing
pipes in congested towns and cities. Furthermore, open trench replacement of Asbestos Cement
(AC) pipes is prohibited by health regulations and recently, bursting of AC pipes in Australia
was also banned (Scott, 2015). Therefore, new low-cost pipe replacement or rehabilitation
techniques are needed for ageing and deteriorated water pipes.
Rehabilitation of aging water pipes by Cured-In-Place-Pipe (CIPP) liners is a relatively new
practice in Australia. CIPP lining technique is a well-established rehabilitation method in which
a resin-saturated tube is introduced into the deteriorated host pipe by either inversion or pull-
in-and-inflate process. After the resin is cured at elevated temperature or using ultraviolet (UV)
light, the resin-impregnated fabric forms a new pipe inside the host pipe. The external surface
of the CIPP liner is in close contact with and conforming to the internal surface of the host
pipe. The internal surface of the CIPP liner is typically a smooth surface which helps reduce
the friction and improve the water flow. The installed CIPP liner can be considered either fully-
structural or semi-structural, depending on the type of liner and its thickness. The main benefits
of using CIPP liners for water pipeline rehabilitation include minimal disturbance to the
community, the ability to line through bends and non-circular shapes (e.g. oval, elliptical) and
the liner itself being free of Volatile Organic Compounds (VOCs) (Marcino and Blate 2015).
Recently, a number of Australian water utilities have begun trialling CIPP liners in their pipe
networks. Due to the increasing age and deterioration of water pipelines, the Australian water
industry is also aiming to standardize the rehabilitation techniques for water pipelines, in
particular for cast iron and AC pipes.
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1.2 History of CIPP liners
The CIPP process was initially developed in the UK in early 1970’s. The first recorded
application of CIPP technology occurred in 1971 in Hackney, East London and involved
relining a 70 m long and 100-year old sewer. This project was undertaken by an engineer named
Eric Wood and supported by entrepreneurs Doug Chick and Brian Chandler. Following the
successful application, a company Insituform (Latin for “form in place”) Pipes and Structures
Ltd was established and the CIPP technology was marketed (Downey 2010). It should be noted
that the pull-in-and-inflate method was used in the first installation and the inversion method
was only available after coated felt was applied in 1973.
The first patent on the CIPP technology was applied by Eric Wood on August 21, 1970 in the
UK, while the first U.S. Patent on the same process was granted to Eric Wood on February 22,
1977. After granting licenses to British contractors to apply this technology for sewer
rehabilitation in England, Insituform further expanded its business in 1976 by granting licenses
to contractors in Europe, Australia and North America. In 1994, the patent for Insituform's
inversion process expired, which resulted in new competition in the CIPP rehabilitation
industry (Rose and Jin, 2006).
Over the years there have been many new variations made to the original patented CIPP
product. Variations exist in resin types, installation methods, curing methods, and tube
construction and only some of these options are applicable for water main rehabilitation (Figure
1.1). Currently, UV curing is only being used in sewer systems. CIPP technology has been
successfully used for rehabilitation of sewers for almost 50 years, but has only been adapted to
use for drinking water mains in the last 18 years. For example, Aqua-Pipe CIPP liner had been
in use in Canada since 2000 and was first used in the U.S. in Illinois in the summer of 2006
(Vose and Loiacono 2007).
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Figure 1.1 Summary of Cured-In-Place-Pipe Liner Technologies (Morrison et al. 2013)
1.3 Current status in Australia
In Australia, due to the predominance of polyethylene and PVC fold-and-form and spirally-
wound linings, the use of CIPP has been limited mainly to sewers, laterals and, more recently,
pipe junctions (Allouche et al. 2014). Many utilities in Australia considered that CIPP is a
valuable technology, but shared the concerns over some issues such as jetting for cleaning in
CIPP-lined pipes, cost effectiveness, and wanting to understand CIPP limitations and risks.
Over the years, CIPP technology has gradually increased its market share in rehabilitating
sewers in Australia. In 2009, Insituform Pacific was awarded a sewer rehabilitation contract to
reline deep sewer mains near Australia’s Parliament House in Canberra. This project involved
the rehabilitation of 450 mm and 750 mm sized sewers along with rehabilitation of connecting
maintenance holes (Trenchless Australasia 2009). By using its next generation iPlus Composite
liner which is reinforced with carbon fibre and/or corrosion resistant fibreglass materials, the
engineers were able to reduce the liner design wall thickness to 15.5 mm, compared to the 28.5
mm liner thickness using the standard CIPP liner. In 2010, Kembla Watertech successfully
rehabilitated 1.8 km of a 600 mm diameter sewer trunk main along the Yuelarbah Management
Trail in Glenrock State Conservation National Park, New South Wales, using a pressure grade
CIPP liner (Trenchless Australasia 2011). This trunk main was part of the Hunter Water
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upgrade management plan for the Dudley Charlestown upgrade project. In 2014, Insituform
Australia was awarded a five-year contract to conduct CIPP lining of wastewater pipelines
ranging from 150–230 mm in diameter for Barwon Water in Victoria, Australia (Trenchless
Australasia 2014).
For water pipelines, Ventia successfully installed Aqua-Pipe®, a structural CIPP liner, to
rehabilitate a 300 mm diameter, cement-lined cast iron and mild steel drinking water main of
Queensland Urban Utilities in 2017. The host pipe was a deteriorated 50-year-old water main
located under a busy railway line. According to Ventia, this installation was the first time that
a structural CIPP liner was used to reline a drinking water pipeline in Australia (Trenchless
Australasia 2017).
2. Liner type classification Lining systems (CIPP and spray liners) used for rehabilitating potable water pipelines can be
classified into four categories by AWWA M28 (2014) and ISO 11295 (2017). A summary of
each liner category obtained from AWWA M28 (2014) and ISO 11295 (2017) is provided
below. Recommended pipe liner class for different modes of failure is shown in Table 2.1.
2.1 Class I
Class I liners are non-structural. The main purpose of a Class I liner is to protect the host pipe
from corrosion, which can improve the hydraulic capacity (reduces build-up of corrosion
products and tubercles) of a structurally sound host pipe. The liner is typically sprayed, and
generally no structural support is expected from the liner. Class I liners require adhesion to the
host pipe. The liner has minimal ability to bridge joint gaps and corrosion holes. In addition, it
is assumed that Class I liners do not contribute to leakage reduction. The internal cement liners
that are installed in many Australian water pipelines since the 1930s can be considered as Class
I liners. The UK has used epoxy resin, polyurea, and polyurethane as Class I liners to improve
water quality and flow; however, these are less commonly used in Australia.
2.2 Class II Class II liners are semi-structural liners that are being used to improve water quality and
hydraulic capacity (varies depending on host pipe condition and liner thickness). Class II liners
require adhesion to the host pipe and expected to extend the life of a partially deteriorated pipe
by reducing leaks and associated accelerated corrosion. All loads are transferred to the stiff
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host pipe (in metallic pipes) and therefore, the liner sustains internal pressure at only
discontinuities in the host pipe (such as corrosion holes). Some common Class II liners used at
present are polyurethane or polyurea type liners.
2.3 Class III Class III liners are similar to Class II liners with the exception that Class III liners should be
able to withstand specified external hydrostatic or vacuum loads (do not rely on adhesion to
the host pipe). Class III liners are CIPP or fibre-reinforced spray liners (fibre reinforced spray
liners are new on the market and of unconfirmed liner class).
2.4 Class IV A Class IV liner is a fully-structural liner and must be able to withstand host pipe conditions
including partially deteriorated, fully deteriorated, reduced ring stiffness, leaks through pipe
barrel or joints, circumferential failures and longitudinal splits. Class IV liners are suitable for
pipes in a deteriorated state (pipes with through-holes, leaks and cracks). Class IV liners should
be tear-resistant and have the ability to hold water pressure under the failure of the host pipe.
Connections, joints and end seals must be adhered or sealed to the liner. The liner does not
need to adhere to the pipe, however water tightness must be satisfied. Typical Class IV liners
are CIPP liners with glass or fibre reinforced layers.
Recommended pipe liner class for different modes of failure is shown in Table 2.1.
Table 2.1. Recommended pipe liner class for different modes of failure in host pipes (adapted
from (AWWA M28 2014; Ellison et al. 2015; ISO 11295 2017).
Estimated future condition of pipe Class I Class II Class III Class IV
Minimal deterioration (no corrosion pits) Yes
Isolated corrosion pits (including through holes) Yes
Leaking joints Yes Yes
Reduced ring stiffness (vacuum, external loads) Yes Yes
Burst failure, circumferential (broken back)
failure, shear failure Yes
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3. CIPP Liner types For potable water pipes, there are four main types of structural CIPP lining systems considered
in this research project:
i. Insituform InsituMain
ii. Sanexen Aqua-Pipe system
iii. Karl Weiss Starline
iv. Sanivar Saniline W
A summary of each liner type is provided in subsequent sub-sections of this report. Note that
the following information are obtained from the respected manufacturers / applicators websites
and no other references were used.
3.1 InsituMain®/Insituform
Figure 3.1 InsituMain® Cross Section (Aegion 2018)
This liner consists of polyethylene-coated, woven glass and polyester fibre lining tube
impregnated with an epoxy resin as shown in Figure 3.1. The coating layer on the inner wall
of the liner serves as a corrosion barrier and to reduce surface friction. The InsituMain® was
introduced as a Class IV fully-structural CIPP liner for pressure pipes following AWWAM28
classification and it is certified to meet the NSF/ANSI Standard 61. The InsituMain® CIPP
liner is applicable to pipe diameters ranging from 150 to 2400 mm. The liner is able to handle
bends up to 45˚ and is pressure-rated above 1.72 MPa. It can be applied to rehabilitate host
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pipes made of different materials, such as cast iron, ductile iron, steel, asbestos cement,
reinforced concrete pipe and thermoplastic pipe and this liner can be installed by either pull-
in-and-inflate or inversion method. After installation, hot water or steam is circulated
throughout the tube to cure the thermosetting resin. After cooling down, the tube ends are cut
off, service connections on the existing host pipe are robotically restored from inside the lined
main (Aegion 2018).
3.2 Aqua-Pipe®/ Sanexen This liner consists of woven textile jacket with epoxy and an inside polymeric membrane for
watertightness as shown Figure 3.2. The Aqua-Pipe® was considered as a Class IV fully-
structural CIPP liner for pressure pipes following AWWA M28 classification and it is certified
to meet the NSF/ANSI Standard 61, UL, BNQ 3660-950, and is a WRAS approved product
(BS6920). The Aqua-Pipe® CIPP liner is applicable to pipe diameters ranging from 150 to 600
mm. According to the manufacturer’s website the liner can possibly be applied through bends,
but no limitations are provided and it is pressure-rated for 1 MPa. It can be applied to
rehabilitate host pipes made of different materials, such as cast iron, steel, asbestos cement and
ductile iron and the liner is installed by pull-in-and-inflate method. The liner is cured by using
hot water, depending on the specific manufacturer's process. The laterals are reconnected by
cutting and reaming the liner with specialized robotic equipment. The liner can be installed up
to 300 m length between access pits (Sanexen Water Inc. 2018).
Figure 3.2 Aqua-Pipe® Cross Section (www.Aqua-Pipe.com)
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3.3 Starline®/Karl Weiss Technologies
Karl Weiss Technologies have developed two different liner products for rehabilitating water
pipelines. Starline® HPL-W is developed for the rehabilitation of transmission and long
distance pipelines while Starline® 1000 is developed for the rehabilitation of distribution
pipelines. Two liner products are certified to meet the NSF/ANSI Standard 61, the UBA
guidelines for drinking water compatibility, and the DVGW worksheet W 270.
The Starline® HPL-W liner consists of a layer of seamless fabric, adhesive and an impermeable
surface layer as shown in Figure 3.3. It has been recommended for water pipes with a diameter
of 100 mm up to 600 mm and with a maximum operating pressure of 4 MPa. A self-advancing
pressure drum on crawler tracks is used push the liner following inversion process. No
additional heating is required for the curing process. It can be applied to pipes with a maximum
length of 600 m with small cleaning access pits at maximum distances of 180 m
Figure 3.3 Starline® HPL-W Cross Section (Starline trenchless technology 2018)
The Starline® 1000-technology is for the rehabilitation of underground drinking water
distribution pipes using fabric lining hose. It has been recommended for maximum operating
pressure of 1 MPa. The product can be applied using a mobile rehabilitation unit independent
of the rehabilitation truck which enables reconditioning in areas not directly accessible to
trucks. This liner is a Class IV liner and is cured with water under high temperature.
9
3.4 Saniline® W/ Sanivar
The liner consists of a layer of polyurethane adhesive, a circular woven jacket of polyester
multi-filament yarn and an internal polyethylene coating as shown in Figure 3.4. The coating
layer on the inner wall of the liner serves as a corrosion barrier and to reduce surface friction.
The liner is certified to meet the requirements of Australian Standard AS/NZS 4020, DVGW
worksheet W 270 and have KTW recommendations for all piping materials. The liner is
applicable to pipe diameters ranging from 100 to 600 mm and is pressure-rated above 1.6 MPa.
It can be applied to rehabilitate host pipes made of different materials, such as cast iron, ductile
iron, steel, asbestos cement pipes and the liner is installed through an inversion process using
a specialist pressurized drum and an average inversion pressure of 1 Bar. It can be applied to
pipes with a workable length of 200 m. The liner is in a soft state when it is installed so it will
conform to the host pipe shape making it suitable for lines with bends and changes in direction
(limitation of maximum bend angle is not provided). Saniline® W is a Class II liner and is cured
with ambient curing methods.
Figure 3.4 Saniline® W Cross Section (www.interflow.com.au)
3.5 Installation of liners
The following steps are required during the installation process:
• Excavation of access pits and the bypassing of the pipe to be rehabilitated;
• Preparation of the water pipe for lining is by cleaning and restoring the cross section.
Cleaning is usually undertaken by high pressure jetting, drag scraping, or rack feed
boring. The pipe surface should be free from debris and running/static water;
• CCTV / careful survey is conducted to accurately locate and plug the lateral connections
to prevent resin from flowing into the lateral connections;
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• CIPP liners are impregnated with resin either in a factory setting or at the site, which is
called the “wet out” process;
• Impregnated liner is inserted into the host pipe to be installed. This can be achieved by
either the pull-in-and-inflate method or the inversion method (Figure 3.5);
• The resin is subsequently cured at ambient or elevated temperature to form a new pipe;
• Pressure tight service connections (Ellison et al. 2015) and any cut ends on the CIPP
liner;
• Service reinstatement can be undertaken externally by access to the lined pipe by local
excavation from the ground surface or internally by locating and reinstating using a
cutter (Robotically);
• Pressure testing for lined pipes is prescribed in ASTM F 1216 (2016).
a) b)
Figure 3.5 CIPP installation methods: a) Pull-in-and-inflate; b) Inversion (Sterling et al.
2010)
4. Liner imperfections, quality standards and installation risks
Refer to “Information on CIPP liner imperfections and common problems.docx”.
5. Design methodology 5.1 Current methodologies in standards
For design of CIPP liners for pressure pipes, there are currently two standards available,
namely, ASTM F1216 (2016) and ASTM F2207 (2013). In ASTM F1216 (2016), the defect is
assumed to be circular while in ASTM F2207 (2013), the defect is considered to be uniquely
11
characterised by two dimensions, w in the hoop direction, and L in the axial direction (Figure
5.1).
Figure 5.1 Defects in a host pipe
Design of CIPP liners is commonly conducted using ASTM F1216 (2016), which was intended
for low pressure force mains and can be used for both partially and fully deteriorated pipes.
For partially-deteriorated pressure pipes, the liner design equation was derived based on the
assumption that the CIPP liner acts like a uniformly pressurised round plate with fixed edges
covering an existing hole in the pipe. Under this assumption, bending stress at and around the
hole controls the design thickness. This is essentially the design for hole-spanning. If this
assumption is not satisfied, the CIPP liner cannot be considered as a circular flat plate and ring
tension or hoop stress will be dominant. In this case, the CIPP liner is considered to be designed
for a fully deteriorated pressure pipe. Apart from internal pressure, the CIPP liner in ASTM
F1216 (2016) is also designed to support hydraulic loads for partially deteriorated host pipes
and to support hydraulic, soil and live loads for fully deteriorated host pipes.
ASTM F2207 (2013) is intended for use in either structurally sound or partially deteriorated
metallic gas pipes. An analytical solution was developed to determine the ultimate strength of
the liner. In the analytical solution, a defect is characterized by dimensions in the hoop and
axial directions respectively. Under internal pressure, the liner is considered to be constrained
to pass through the end-points of the defect, as it bulges out of the defect. In terms of material
properties, the liner is assumed to be an orthotropic membrane without any shear stiffness. The
shape of the liner in the axial and hoop directions is also assumed to be a circular arc. Two
failure criteria, namely the maximum stress and interactive criteria, can be used to calculate the
burst pressure of the liner.
Circular defects Rectangular defects
𝑤𝑤 𝐿𝐿
𝑤𝑤
𝐿𝐿
12
In ASTM F2207 (2013), a number of mechanical tests are suggested to be conducted to
determine the mechanical behaviour of the liner. The tests include peel tests, strength tests, and
flexibility tests. The strength tests may be the sustained pressure test of a lined deteriorated
metallic pipe with a full circumferential gap between two pipe segments and a hole in the host
pipe, using an extension of either the test method ASTM D1598 (2002) or the test method
ASTM D2837 (2011). The flexibility tests include tensile tests and bend tests of a lined pipe
with a ring fracture in the middle of the host pipe subjected to the maximum allowable
operating pressure.
5.2 Gap identification
For partially deteriorated host pipes, the design equation for pressure pipes in the current design
standard ASTM F1216 (2016), only accommodates a circular hole in the host pipe and isotropic
liner materials while the design equation in ASTM F2207 (2013) considers a rectangular
defect, characterised by two dimensions in the hoop and axial directions, respectively and
anisotropic liner materials. It can be seen that current design standards/guidelines only support
either circular or rectangular holes in the host pipe under internal pressure only. Both standards
are not able to consider the effect of an inclined rectangular/elliptical defect, a longitudinal
crack-like defect, a ring fracture/damaged joint.
In terms of external loads, although hydraulic, soil and live loads are considered in ASTM
F1216 (2016), pressure transients, pressure induced thrust forces, Poisson effect due to the use
of anisotropic liner materials, thermal expansion effects and differential ground movement
have not been taken into account in both standards. In addition, the effects of liner
imperfections were not considered in both standards. For a future standard, these effects need
to be examined.
6. Review on Experimental Testing 6.1 Short-term tests
Various short-term tests have been performed to investigate the behaviour of the CIPP liners.
These tests can in general be classified into specimen tests and large-scale pipe tests.
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6.1.1 Specimen tests
6.1.1.1 Tensile tests
Tensile tests were conducted by Allouche et al. (2005) on Aqua-Pipe liners. Uniaxial test
specimens cut in the longitudinal direction of the liner were prepared by removing the liner
specimens from cast iron host-pipes. Due to specimen curvature across the short direction, end
tabs of epoxy resin were used to achieve effective gripping. The measured stress strain curve
(Figure 6.1) showed that the behaviour of the material in short term can be represented by a
bilinear curve, with Young’s modulus of 2 GPa up to a strain of 1.3%, and a subsequent
modulus of 180 MPa.
Figure 6.1 Uniaxial stress-strain behaviour; test data and bilinear fit (Allouche et al. 2005)
Brown et al. (2008) investigated the mechanical properties of a composite liner in both the
longitudinal and circumferential directions at different temperatures. The experimental
methodology and test specimen configuration were based on ASTM D3039/D3039M (2006).
The testing results were summarised in Table 6.1 and Table 6.2. It was found that the initial
modulus and ultimate strength in the hoop direction were about 45% higher than those in the
longitudinal direction. It was also found that at higher curing temperature, the liner strength
and stiffness were higher, up to a temperature of 55˚C.
𝐸𝐸1 = 2 𝐺𝐺𝐺𝐺𝐺𝐺
𝐸𝐸2 = 180 𝑀𝑀𝐺𝐺𝐺𝐺
𝜀𝜀𝑦𝑦 = 1.3%
14
Table 6.1 Effect of curing temperature on the modulus and strength of the resin (Brown et al.
2008)
Curing
temperature
(˚C)
Tensile modulus (MPa) Tensile strength (MPa)
No. of
specimens Mean ± STD
No. of
specimens Mean ± STD
20 5 1832 ± 162 3 47.5 ± 0.6
40 5 2076 ± 90 2 52.7 ± 0.8
55 5 2356 ± 133 2 60.8 ± 0.6
70 4 2307 ± 117 0 -
Table 6.2 Tensile properties of the exhumed liner and fabricated liner (Brown et al. 2008)
Liner
type
Sample
orientation
No. of
specimens
Tensile
modulus
(MPa)
Yield
strength
(MPa)
Yield
strain
(%)
Ultimate
tensile
strength
(MPa)
Exhumed Longitudinal 4 2019 ± 8.6 23.5 1.0 61 ± 0.6
Fabricated Longitudinal 5 2017 ± 243 24 0.9 61.3 ± 2.8
Circumferential 5 3040 ± 120 24 0.9 88.4 ± 4.7
Interplastic Corporation (2008) examined the differences in tensile properties between
laboratory-prepared and field-obtained CIPP sewer liner samples. The resin/felt composites
were constructed by impregnating 6 mm, needle-punched, polyester fabric felt with an
applicable resin/initiator system. Static tensile properties of the specimens were tested
according to ASTM D638 (2008). The testing results were summarised in Table 6.3. Results
showed that the tensile strengths and tensile moduli of the laboratory samples are marginally
greater than the 10 samples obtained from the field.
Matthew et al. (2012c) determined the tensile properties of the retrieved liner samples (Aqua-
Pipe) from a 152 mm diameter cast iron water pipe installed in 1914 and 1949 in the city of
Cleveland. A total of five specimens, which were cut in longitudinal direction, were prepared
and tested in accordance with ASTM D638 (2008). The stress and strain curves of the tests
were presented in Figure 6.2. It was found that the average of the tensile strength was 65 MPa
with a standard deviation of 2.1 MPa while the average of the tensile modulus was 3559 MPa
with a standard deviation of 1054 MPa.
15
Table 6.3 Comparison of tensile properties of laboratory and field-generated samples
(Interplastic Corporation 2008)
Sample ID
Sample
Acquisition
Source
Resin
Content %
Tensile
Strength
(MPa)
Tensile
Modulus
(MPa)
Degree of
Cure (%)
F-1 Field 78.9 21.2 4780 99+
F-2 Field 78.9 21.9 4050 95.5
F-3 Field 79.7 23.6 4690 97.6
F-4 Field 79.54 22.9 4570 99+
F-5 Field 77.81 21 4540 98.2
F-6 Field 77.94 21.9 4410 99+
F-7 Field 80.47 20.8 4300 97.2
F-8 Field 79.82 20.8 4250 99+
F-9 Field 78.72 23.2 4620 99+
F-10 Field 78.98 22 4610 97.8
Average/STD 21.93/1.02 4482/225
L-1 Laboratory 85.66 24.1 4580 99+
L-2 Laboratory 70.31 29 4450 99+
L-3 Laboratory 66.09 27.6 4620 99+
Average/STD 26.9/2.52 4550/89
Figure 6.2 Stress-strain Curves from Tensile Testing (Adapted from Matthew et al. 2012c)
0
20
40
60
80
0 0.02 0.04 0.06 0.08 0.1 0.12
Stre
ss (M
Pa)
Strain
Tensile stress vs. strain
Sample1Sample2Sample3Sample4
16
Allouche et al. (2012) and Allouche et al. (2014) conducted tensile testing on old CIPP liners
for sewer pipes in the city of Denver and the City of Columbus. For all tests, specimens were
prepared following ASTM D638 (2008) standard. The city of Denver had two test sites. At Site
1, the tested liner was a 25-year old Insituform CIPP liner (Unwoven fabric with polyurethane)
in a 203 mm diameter clay pipe. The outer diameter and thickness of the liner were 203 mm
and 6 mm respectively. Tests were conducted by both the Trenchless Technology Centre (TTC)
and Insituform. A total of nine specimens were prepared and tested (three from each of the
crown, spring line and invert locations). At Site 2, the tested liner was a 23-year old Insituform
CIPP liner installed in a 1219 mm diameter brick sewer pipe. The liner thickness was 18 mm
upstream and 13.5 mm downstream. One pipe sample taken from the CIPP lined pipe was
removed and tested by Insituform in 1995. A total of five specimens were prepared and tested
for the downstream and the upstream, respectively. The City of Columbus also had two test
sites. At Site 1, the liner was a 5-year old Reynolds Inliner® (now known as Layne inliner)
CIPP liner installed in a 203 mm diameter clay pipe. At Site 2, the liner was a 21-year old
Insituform CIPP liner installed in a 914 mm diameter brick sewer. For each site, a total of 15
specimens were prepared and tested (five specimens from each of the crown, spring line and
invert locations). The tensile testing results were summarised in Table 6.4.
Table 6.4 Tensile properties for old sewer CIPP liners (Allouche et al. 2012)
City Host
Pipe
Diameter
(mm)
Liner
Age Liner Set
Tensile Strength
(MPa)
Tensile Modulus
(MPa)
Denver
Clay
pipe 203 25 Insituform
By TTC 21±1.2 2838±280
By
Insituform 16±1.4 -
Brick
sewer 1219
23 Insituform
Upstream 1 22±1.5 2943±403
Upstream 2
Downstream 21±1.6 2636±415
8 16±1.2
Columbus
Clay
pipe 203
5 Reynolds
Inliner
27±2.9 2500±301
0 - -
Brick
sewer 914 21 Insituform 20±1.7 2174±293
A demonstration project for the installation of a UV cured CIPP liner (Reline America Blue-
Tek™ liner, now known as Alphaliner) was conducted in the city of Frisco, Texas (Matthews
2014). A 271 m section of 250 mm diameter vitrified clay pipe (VCP) was lined with UV cured
17
CIPP liner. For comparison with the field-obtained samples, an 18 m long PVC pipe with 250
mm diameter was lined above ground to provide extra liner samples under controlled
conditions. The tensile tests of the specimens obtained from the retrieved liner samples were
performed by Matthews (2014) based on ASTM D638 (2008). Six samples were cut from the
field lined VCP pipe section while one was cut from the lined PVC pipe. A total of five test
specimens cut in the longitudinal direction were prepared and tested for each of the seven
samples. The testing results were summarised in Table 6.5. It was found that tensile strength
results of the above ground samples are higher than those of the field-obtained liner samples.
The variations in both tensile strength and tensile modulus of the field-obtained liner samples
are quite high (>20%).
Table 6.5 Tensile properties for UV cured CIPP liners (Matthews 2014) Sample set No. of Samples Tensile Strength (MPa) Tensile Modulus (MPa)
Above ground sample 5 166 ± 21 12500 ± 2000
Field set 1 6 147 ± 27 14000 ± 3000
Field set 2 5 144 ± 31 17800 ± 8000
Field set 3 6 158 ± 26 12600 ± 2400
Field set 4 5 143 ± 24 13600 ± 3000
Field set 5 7 138 ± 26 16800 ± 3800
Field set 6 5 137 ± 19 10100 ± 2400
Average 39 147 ± 25 14000 ± 4400
Cornell University tested two sections of CIPP lined 150-mm-diameter cast iron (CI) gas pipe
and two sections of lined 300-mm-diameter CI gas pipe with joints (Stewart et al. 2015). All
the four sections were approximately 2.4 m long with the joint located at the centre of the
section. Two sets of specimens were considered for the 150 mm-diameter lined CI pipe. One
set were specimens which had experienced aging in field for 16 years while the other set were
specimens that had experienced aging in field for 16 years together with mechanical aging
equivalent to 100 years. Similarly, two set of specimens were considered for the 300-mm-
diameter lined CI pipe. One set were specimens which had experienced aging in field for 10
years while the other set were specimens that had experience aging in field for 10 years together
with mechanical aging equivalent to 100 years.
Tensile tests were conducted on bonded and de-bonded CIPP liner specimens for both the 150
mm and 300 mm-diameter CI pipes, as show in Figure 6.3. The tests followed the modified
ASTM D3039/3039M (2000), based on the testing by Netravali et al. (2003) for investigating
the behaviour of Starline®2000 PSE-35 liner. The liner specimens were tested in both
18
longitudinal and circumferential directions, with 15 mm in width and 200 mm in length. The
thickness values of the liner for 150-mm-diameter pipe and 300-mm-diameter pipe are 1.25
mm and 1.82 mm respectively. The testing results were summarised in Table 6.6. Results
showed a large difference in the circumferential and longitudinal tensile strength
(circumferential = 23 - 46.7 MPa, longitudinal = 76.9 - 137.5 MPa) of Starline®2000 PSE-35
liner.
Figure 6.3 Tension test (Stewart et al. 2015)
Table 6.6 Tensile properties of Starline CIPP liner (Stewart et al. 2015)
Host
pipe
Diameter
(mm) Liner Orientation Set Set type
Tensile strength
(MPa)
Secant
modulus
(MPa)
Cast
Iron
150
Starline®
2000 PSE-
35 liner
Longitudinal
Field aged (16
years) 131.2±9.7
760
Field (16
years) and
Mechanically
(Equi. to 100
years) aged
De-bonded
1 137.5±7.5
De-bonded
2 126.3±7.1
Bonded 1 125.1 ± 4.4
Bonded 2 118.8 ± 3.5
Circumferential
Field aged (16
years) 24.8 ± 1.2
Field (16
years) and
De-bonded
1 24 ± 1
19
Mechanically
(Equi. to 100
years) aged
De-bonded
2 23 ± 0.8
Bonded 1 23 ± 0.9
Bonded 2 24.5 ± 0.9
300
Longitudinal
Field aged (10
years) 80.5 ± 6.3
Field (10
years) and
Mechanically
(Equi. to 100
years) aged
De-bonded
1 79.6 ± 5.1
De-bonded
2 80.7 ± 6.3
Bonded 1 87.2 ± 1.6
Bonded 2 82.9 ± 7.6
Circumferential
Field aged (10
years) 40.8 + 2.4
Field (10
years) and
Mechanically
(Equi. to 100
years) aged
De-bonded
1 44.6 ± 2.5
De-bonded
2 45.2 ± 1.9
Bonded 1 46.7 ± 2.9
Bonded 2 44.1 ± 2.6
Sterling et al. (2016) conducted a retrospective study on CIPP liners installed in gravity sewers.
18 CIPP liner samples were collected (aged from 17 to 34 years), while two younger liners (5
and 9 years) were also included. The tensile testing followed ASTM D638 (2014). The testing
results were summarised in Table 6.7. For the 18 samples, the mean and standard deviation of
the tensile strength are 23 and 3 MPa respectively. The mean and standard deviation of the
tensile modulus are 2851 and 455 MPa respectively (Note: tensile testing was only conducted
in the longitudinal direction).
Yan (2016) conducted tensile testing for a CIPP liner (a new fibre-reinforced composite hose)
manufactured by Asoc from China for gas pipelines. The testing followed ASTM D638 (2014)
with type 1 specimen shape. A total of fifteen specimens with an average thickness of 6.2 mm
were prepared, with five each in the longitudinal, circumferential, and inclined (45 degrees
from the pipe axis) directions, respectively. The testing results are summarized in the following
table.
20
Table 6.7 Measured tensile properties of retrospective samples (Sterling et al. 2016) Sample retrieve location Diameter of the pipe
(mm)
Age (years) Average values
Tensile – ASTM D638 (MPa)
Strength Modulus
Columbus 900 21 20 2174
Columbus 200 5 27 2500
Denver 200 25 21 2838
Denver Downstream 1200 23 21 2637
Denver Upstream 1200 23 22 2943
Edmonton 250 19 22 3011
Edmonton 200 19 25 3517
Houston 525 17 24 3208
Houston 450 17 22 3110
Indianapolis 1050 25 19 2422
Nashville Dunston 200 19 24 2591
Nashville Wyoming 200 9 18 2764
New York City 375 23 26 3821
New York City 300 24 23 2237
Northbrook 300 34 30 2989
Winnipeg Richard 750 34
Winnipeg Kingsway 450 34
Winnipeg Mission 750 28
Average 23 2851
Standard deviation 3 455
Percent Standard deviation 13.7 16.0
Table 6.8 Tensile testing results (Yan 2016) Direction Tensile strength (MPa) Tensile Modulus (MPa)
Longitudinal Average 31.9 470.1
STD 7.7 43.5
Circumferential Average 25.4 178.7
STD 10.3 70.3
Inclined (45 degrees from
the pipe axis
Average 4.0 67.2
STD 0.9 31.9
6.1.1.2 Flexural tests
Lystbaek (2007) investigated the field performance of CIPP liners installed in Aarhus,
Denmark. A total of six pipe samples were collected for testing. All the liners from which the
samples were taken were installed between 1991 and 1992. The sampling was conducted in
1999/2000 and 2005. Flexural testing was performed according to ISO 178 (2010) and the
21
results were summarised in Table 6.9. No significant difference in long-term strength and
modulus were observed after years in service.
Table 6.9 Three-point flexural test data for the retrospective liner sampling study in Denmark
(Lystbaek 2007)
Liner
diameter
(mm)
Wall
thickness
(mm)
Flexural modulus (MPa) Flexural strength (MPa)
1991/1992 1999/2000 2005 1991/1992 1999/2000 2005
Pipe 1 200 6 2615 2623 3870 39 40 37.8
Pipe 2 200 6 2608 2903 3147 37 37 35.1
Pipe 3 200 6 2608 3160 3274 37 35 39.6
Pipe 4 400 9 2396 4120 3463 40 49 45.8
Pipe 5 250 6 2790 3735 3433 43 44 44.3
Pipe 6 500 9 - - 3697 - - 46.4
Average/STD 2603/140 3308/612 3480/266 39.2/2.5 41.0/5.6 41.5/4.7
Interplastic Corporation (2008) examined the differences in flexural properties between
laboratory-prepared and field-obtained CIPP liner samples. Static flexural properties of the
liner were tested according to ASTM D790 (2007). The testing results were summarised in
Table 6.10. Results showed that the flexural strength and flexural moduli of the laboratory-
prepared samples are relatively higher than those of the field-obtained samples. In addition, it
was observed that the flexural properties are not influenced by the percentage of resin system
in the composite and the surface quality has a major effect on the flexural strength but a minor
effect on the flexural modulus.
Table 6.10 Comparison of flexural properties of laboratory and field-generated samples
(Interplastic Corporation 2008)
Sample ID
Sample
Acquisition
Source
Resin Content
%
Flexural
Strength
(MPa)
Flexural
Modulus
(MPa)
Degree of
Cure %
F-1 Field 78.90 43.6 3880 99+
F-2 Field 79.94 50.4 4000 95.5
F-3 Field 79.70 49.7 3670 97.6
F-4 Field 79.54 42.4 3750 99+
F-5 Field 77.81 43.3 3940 98.2
F-6 Field 77.94 49.3 3920 99+
F-7 Field 80.47 37.6 3570 97.2
F-8 Field 79.82 47.3 3700 99+
22
F-9 Field 78.72 45.3 3830 99+
F-10 Field 78.98 46.9 3700 97.8
Average/STD 45.6/3.97 3796/139
L-1 Laboratory 85.66 65.2 4850 99+
L-2 Laboratory 70.31 52.5 3590 99+
L-3 Laboratory 66.09 70.9 4280 99+
Average/STD 62.9/9.42 4240/631
Matthew et al. (2012c) investigated the flexural properties of the retrieved Aqua-Pipe liner
samples from a relined 152 mm diameter cast iron pipe in the city of Cleveland. Five
specimens, cut in longitudinal direction of the liner, were tested according to ASTM D790
(2007). The stress strain curves were presented in Figure 6.4. It was found that the average
flexural strength was 55 MPa with a standard deviation of 4 MPa while the average flexural
modulus was 2530 MPa with a standard deviation of 105 MPa.
Figure 6.4 Stress vs. Strain Curves for flexural testing (Adapted from Matthew et a. 2012c)
Allouche et al. (2012) and Allouche et al. (2014) also conducted flexural testing on old
Insituform and Reynolds Inliner CIPP liners for sewer pipes in the city of Denver and the city
of Columbus. Specimens were cut from the retrieved CIPP liner and tested based on ASTM
D790 (2007). The flexural testing results for the four sites were summarised in Table 6.11:
0
10
20
30
40
50
60
0 0.01 0.02 0.03 0.04 0.05 0.06
Flex
ural
stre
ss (M
Pa)
Strain
Flexural stress vs. strain
Sample 1
Sample 2
Sample 3
Sample 4
Sample 5
23
Table 6.11 Flexural properties for old sewer CIPP liners (Allouche et al. 2012)
City Host
Pipe
Diameter
(mm)
Liner
Age Liner
Specimen
batch
Flexural
Strength (MPa)
Flexural
Modulus (MPa)
Denver
Clay
pipe 203 25 Insituform
By TTC 47 ± 3.8 2313 ± 125
By Insituform 48 ± 3.1 3406 ± 295
Brick
sewer 1219
23 Insituform
Upstream 1 34.7 ± 4.5 1259 ± 159
Upstream 2 42 ± 6.1 1818 ± 485
Downstream 48.5 ± 2.4 2089 ± 168
8 48 ± 2.8 3378 ± 276
Columbus
Clay
pipe 203
5 Reynolds
Inliner
44 ± 14.1 2386 ± 343
0 50 ± 3.4 -
Brick
sewer 914 21 Insituform 42 ± 2.7 1426 ± 200
Another large-diameter CIPP liner demonstration (Matthews 2014) was conducted for a water-
cured CIPP liner. The demonstrated pipe was part of the Elm Fork Relief Interceptor system,
owned by Trinity River Authority of Texas in Irving, Texas. A 238 m section of a 2400 mm
diameter reinforced concrete pipe (RCP) was lined with water-cured CIPP lining. Flexural
strength and modulus tests, following ASTM D790 (2007) and ASTM F2019 (2011),
respectively, were conducted on a total of 72 specimens obtained from the retrieved liner
samples. The testing results were summarised in Table 6.12. It was found that the measured
flexural properties exceeded the design and suggested specification from the manufacturer. It
was also emphasised in this study that fibreglass liners must be tested in accordance with
ASTM F2019 (2011), which requires a 50 mm wide specimen cut from the circumferential
direction.
Table 6.12 Flexural properties for UV and Water cured CIPP liners (Matthew 2014) Location Host pipe Diameter
(mm)
Liner Orientation Set No. of
samples
Flexural
Strength
(MPa)
Flexural
Modulus
(MPa)
City of
Frisco
vitrified
clay pipe
250 UV cured
CIPP liner
(Reline
America
Blue-Tek™
liner)
Longitudinal Above
ground
sample
5 253 ± 50 15100 ± 1900
Field set 1 6 210 ± 54 9700 ± 2600
Field set 2 5 163 ± 58 9500 ± 3900
Field set 3 6 153 ± 68 7500 ± 3500
Field set 4 5 218 ± 93 10800 ± 4900
Field set 5 7 224 ± 48 11200 ± 2600
Field set 6 5 175 ± 61 7900 ± 3600
Average 39 200 ± 66 10200 ± 3900
24
Circumferential Above
ground
sample
5 506 ± 24 16200 ± 1000
Field set 1 5 321 ± 101 11100 ± 2600
Field set 2 5 307 ± 69 14500 ± 3400
Field set 3 5 475 ± 90 16700 ± 3400
Field set 4 5 374 ± 88 10200 ± 2700
Field set 5 5 386 ± 61 11500 ± 2000
Field set 6 5 390 ± 51 11700 ± 2800
Average 35 390 ± 93 13100 ± 3400
Irving
reinforced
concrete
pipe
2400 Water-
Cured CIPP
liner
(Insituform
iPlus®
Composite
Liner)
- Field set 1 5 89 ± 8 6900 ± 200
Field set 2 5 96 ± 3 7400 ± 300
Field set 3 5 84 ± 1 6700 ± 200
Field set 4 5 84 ± 1 6900 ± 100
Field set 5 5 92 ± 3 7200 ± 200
Field set 6 5 94 ± 2 7200 ± 100
Field set 7 5 74 ± 2 7000 ± 200
Field set 8 5 69 ± 1 6600 ± 100
Field set 9 5 72 ± 4 6400 ± 200
Field set 10 5 77 ± 5 6700 ± 200
Field set 11 5 73 ± 3 6700 ± 100
Field set 12 5 73 ± 1 7000 ± 100
Field set 13 5 81 ± 2 7000 ± 200
Field set 14 5 73 ± 1 6300 ± 100
Field set 15 5 78 ± 3 7100 ± 100
Average 75 80 ± 9 6900 ± 300
In the retrospective study conducted by Sterling et al. (2016) on CIPP liners for gravity sewers,
flexural specimens were tested according to ASTM D790 (2007). The testing results were
summarised in Table 6.13. For the 17 sites, the flexural strength values ranged from 30.8 to
59.2 MPa while the flexural modulus ranged from 1426 to 3293 MPa. It was stated that it was
not possible to tell whether the low values of the flexural properties was caused by the ongoing
deterioration or the poor liner properties that had existed since installation. The mean and
standard deviation of the flexural strength are 45.4 and 7.4 MPa respectively. The mean and
standard deviation of the flexural modulus are 2189 and 484 MPa respectively.
Table 6.13 Measured flexural properties of retrospective samples (Sterling et al. 2016) Sample retrieve
location
Diameter of the pipe
(mm)
Age (years) Average values
Flexure – ASTM D790 (MPA)
Strength Modulus
Columbus 900 21 42 1426
Columbus 200 5 44 2386
Denver 200 25 47 2312
Denver Downstream 1200 23 48 2089
25
Denver Upstream 1200 23 38 1539
Edmonton 250 19 42 2285
Edmonton 200 19 47 2515
Houston 525 17 48 2328
Houston 450 17 50 2334
Indianapolis 1050 25 32 1636
Nashville Dunston 200 19 47 2080
Nashville Wyoming 200 9 38 1948
New York City 375 23 55 3293
New York City 300 24 50 1966
Northbrook 300 34 54 2223
Winnipeg Richard 750 34 59 3117
Winnipeg Kingsway 450 34 47 2233
Winnipeg Mission 750 28 31 1694
Average 45 2189
Standard deviation 7 484
Percent Standard deviation 16.2 22.1
Shaded boxes indicate data that do not meet the current minimum ASTM requirement.
Yan (2016) also conducted flexual testing for the CIPP liner (a new fibre-reinforced composite
hose) manufactured by Asoc from China for gas pipelines. The testing followed ASTM D790
(2015). Five specimens with an average thickness of 6.2 mm were prepared along the
longitudinal direction. The average flexural strength and flexural modulus are 4 MPa and 25
MPa respectively.
6.1.1.3 Bond Tests Lap shear tests (Figure 6.5) were performed by Stewart et al. (2015) on a Starline®2000 liner.
The tests, in which CIPP liner is pulled from host pipe to determine the shear strength between
the liner and CI host pipe, followed a modified ASTM D3164 (1997) procedure based on the
testing conducted by Netravali et al. (2003). The specimens were cut and tested in the
longitudinal direction of the lined CI pipe. The width and length of the liner specimens were
25.4 mm and 152.4 mm, respectively. The overlap length between the liner and the host pipe
was 8 mm. In the tests, one end of the testing machine gripped the host pipe while the other
end gripped the liner. The testing results were summarised in
Table 6.14.
26
Figure 6.5 Lap shear test (Stewart et al. 2015)
Table 6.14 Lap shear strengths of Starline CIPP liner (Stewart et al. 2015) Host
pipe
Diameter
(mm) Liner Orientation Set Set type
Shear strength
(MPa)
Cast
Iron
150
Starline®
2000 PSE-35
liner
Longitudinal
Control (Lab
prepared samples) 7.6 ± 1
Specimen aged
for 48 weeks at
65 ˚C (Equi. To
22 years)
9.7 ± 0.55
Field aged (16
years) 8.1 ± 0.96
Field (16 years)
and Mechanically
(Equi. to 100
years) aged
Bonded 1 9.2 ± 0.97
Bonded 1 7.7 ± 2.26
300 Longitudinal
Control (Lab
prepared samples) 7.6 ± 1
Specimen aged
for 48 weeks at
65 ˚C (Equi. To
22 years)
11 ± 1.11
Field aged (10
years) 8.8 ± 0.39
Field (10 years)
and Mechanically
(Equi. to 100
years) aged
Bonded 1 8.1 ± 1.69
Bonded 1 10.3 ± 0.76
27
Stetter et al. (2017) designed an experiment to investigate the adhesion of Aqua-Pipe and
Starline STRUCTURE-W liners to the AC host pipe. Specimens were taken from 20 mm
diameter circular sections of the lined pipe (Figure 6.6). Delamination of the liner from the AC
pipe occurred for both liner types during cutting. The weak adhesion between the liners and
the AC host pipe was attributed to the internal surface cleanliness and internal corrosion of AC
host pipe.
Figure 6.6 Lined AC pipe with adhesion samples removed (Stetter et al. 2017)
Figure 6.7 Peel test (Stewart et al. 2015)
Stewart et al. (2015) employed a modified ASTM D1876 (1995) procedure using a 180° peel
test method to determine the peel strength between the liner and the CI pipe (Figure 6.7), based
28
on specimen dimensions used by Netravali et al. (2003). The specimens were prepared along
the direction of the pipe longitudinal axis. The width and length of the specimens were 25.4
mm and 300 mm, respectively. Two thickness values, 1.25 mm and 1.82 mm, of the liner, were
considered. The testing results were presented in the Table 6.15.
Table 6.15 Peel strengths of Starline CIPP liner (Stewart et al. 2015) Host
pipe
Diameter
(mm) Liner Orientation Set Set type
Shear strength
(MPa)
Cast
Iron
150
Starline®
2000 PSE-35
liner
Longitudinal
Control (Lab
prepared samples) 1.4 ± 0.11
Specimen aged
for 48 weeks at
65 ˚C (Equi. To
22 years)
1.7 ± 0.17
Field aged (16
years) 1.53 ± 0.2
Field (16 years)
and Mechanically
(Equi. to 100
years) aged
De-Bonded 1 1.41 ± 0.3
De-Bonded 1 1.02 ± 0.31
300 Longitudinal
Control (Lab
prepared samples) 1.4 ± 0.11
Specimen aged
for 48 weeks at
65 ˚C (Equi. To
22 years)
1.51 ± 0.15
Field aged (10
years) 0.82 ± 0.26
Field (10 years)
and Mechanically
(Equi. to 100
years) aged
De-Bonded 1 0.68 ± 0.2
De-Bonded 1 0.27 ± 0.07
6.1.1.4 Parallel plate loading/Pipe ring tests
Parallel plate loading tests were conducted by Allouche et al. (2005) on short segments of the
Aqua-Pipe liners (4 specimens in total), following ASTM D2412 (2002). It was found that the
initial modulus of 1500 MPa was lower than the longitudinal uniaxial modulus. Due to non-
uniform strains produced by the parallel plate, the modulus decreased gradually to 600 MPa.
29
Softening started at the extreme fibres at the spring lines, crown and invert, followed by
continuous softening of the specimen as strains increased.
6.1.1.5 Split-disk tests
Ampiah et al. (2008) and Ampiah et al (2010) conducted a series of laboratory experiments to
investigate the effect of liner folds on the strength of the Aqua-Pipe liner (Figure 6.8 and Figure
6.9). The testing method followed the split-disk procedure defined in ASTM D2290 (2004).
The dimensions of testing specimens and the testing results were summarised in Table 6.16. It
was observed during the testing that failure initiated at the fold area, which indicated that the
fold is a source of weakness in the liner. The amplitude, angle and size of the fold were found
to significantly affect the load at which the resin in the fold started to crack. However, the final
failure load of the folded specimens decreased only when fold angle and width were large
enough, e.g., width larger than 15 mm. In addition, the adverse effect of the fold can be
minimised when two jackets are in close proximity.
Figure 6.8 Setup for split-disk test (Ampiah et al. 2008)
30
Figure 6.9 Labelled picture of fold types (a) SW; (b) IW; (c) LW; and (d) geometric
parameters (Ampiah et al. 2008)
Table 6.16 Effect of the presence of a fold and its geometry on hoop strength (Ampiah et al.
2008) Sample
ID
fold type Loading rate
(mm/min)
𝚫𝚫
(mm)
𝝀𝝀
(mm)
𝑫𝑫𝒐𝒐
(mm)
𝑷𝑷𝒄𝒄𝒄𝒄
(N)
Average 𝑷𝑷𝒄𝒄𝒄𝒄
(N)
𝑷𝑷𝒎𝒎𝒎𝒎𝒎𝒎
(N)
Average
𝑷𝑷𝒎𝒎𝒎𝒎𝒎𝒎 (N)
NW-1 None 5 0 0 158 N/A N/A 13108 11445
NW-2 0 0 156 N/A 10161
NW-3 0 0 155 N/A 11935
NW-4 0 0 157 N/A 11465
NW-5 0 0 160 N/A 10554
SW-1 Inner jacket
only
5 10.00 12.40 153 6039 6281 11971 11701
SW-2 9.99 11.94 155 6432 12249
SW-3 9.97 10.69 154 6372 10761
SW-4 10.09 10.27 153 N/A 12349
SW-5 10.07 10.71 154 N/A 11176
IW-1 Both jackets 5 16.38 21.96 150 3143 3564 10292 10777
IW-2 16.31 22.65 151 4024 11604
IW-3 17.04 23.09 150 2489 12003
IW-4 15.28 21.60 152 4399 8929
IW-5 15.29 20.24 150 3763 11057
LW-1 Both jackets 5 16.03 26.15 155 6391 5401 8573 8773
LW-2 15.92 25.47 156 3311 8491
LW-3 15.81 32.79 158 4787 8546
LW-4 15.82 - 155 6329 8627
LW-5 15.29 30.09 155 6189 9626
𝐷𝐷𝑜𝑜 is the outer diameter of liner
31
6.1.2 Large-scale pipe tests
6.1.2.1 Pipe pressure tests
Sanivar AG (2000) performed three pressure tests on Saniline W liners. In the first pressure
test, the host pipe was a DN 150 steel pipe of 1.6 m in length. Both ends of the pipe were sealed
properly with flanges (DN 150, PN 10). The steel pipe was prepared with three holes of DN 32
to simulate corrosion damage in the old pipe. In the testing process, the internal pressure was
increased from 1.6 to 4.8 MPa and it was found that the liner maintained its pressure integrity.
The second test was a pressure test of stiffness and elasticity of the liner at pipe joints. Two
pieces of a DN 300 AC pipe with a wall thickness of 25 mm and a length of 0.6 m each were
lined with Saniline liners. A gap of 25 mm between the two pipe pieces was used to simulate
a joint. At both ends, additional DN 300mm steel pipes were used to further extend the AC
pipes and welded struts were used for reinforcement. The testing pressure was increased to
2.85 MPa and no leaking of water though the Saniline W liner was found. The third test was a
bust test. The host pipe was a DN 300 steel pipe of 1.7 m in length. A DN 100 hole was drilled
in the host pipe to simulate potential corrosion damage. Welded struts were used for
reinforcement at both ends of the pipe. When the internal pressure was increased to 2.6 MPa,
it was observed that the holding devices were about to deform and therefore the test was
stopped.
Allouche et al. (2005) and Allouche and Moore (2005) developed a pressure testing program,
including seven burst tests on a 75 year old 150 mm diameter cast iron water main lined with
Aqua-Pipe liners. All the tests were performed using a custom-made testing apparatus, which
is capable of producing an internal water pressure of up to 5 MPa. Both ends of the specimens
were first squared using a lathe for sealing purposes and then treated with liquid rubber. 12
high-yield threaded bars were employed to connect the steel end bulkheads. A compression
force was applied to the CI pipe in order to resist internal pressure blowout. The set-up of the
testing facility was shown in Figure 6.10. It should be noted that the testing set-up can resist an
internal pressure of as high as 3.8 MPa for short durations, even across large gaps in the host
pipe. During the testing, some permanent liner deformation was noted under a constant pressure
of 1.3 MPa over 500 hours. The test was stopped at a pressure of 3.5 MPa when fibre cracking
in the liner occurred.
32
Figure 6.10 Test setup for pressure testing (Allouche et al. 2005)
Matthews et al. (2012c) performed a vacuum test on Aqua-Pipe CIPP liners lined on a ductile
iron host pipe (Figure 6.11). The CIPP liner was first carefully removed from the host pipe
using a manually operated hydraulic press while minimising the possibility of damaging the
liner. Then a 0.6 m long section was cut out from the removed liner using a hand saw. The two
end caps for sealing the liner specimen were manufactured by cutting a circular steel plate of
165 mm in diameter, and welding it to a short segment of a steel pipe (76 mm long). A provision
to a quick connector was attached to one cap and polyurea was next poured inside the cap. The
liner was then inserted into the cap and held in position until cured. The procedure was repeated
for the other end of the specimen. Testing results showed that the liner withstood the vacuum
force (-0.1 MPa) for 70 hours with small deflection (up to 0.06 mm).
Figure 6.11 Complete Experimental Setup (Matthews et al. 2012c)
33
6.1.2.2 Pipe bending tests
Allouche and Alam (2012) investigated the behaviour of Aqua-Pipe Liner subjected to bending
(Figure 6.12) under both pressurised and non-pressurised conditions at the location of ring
fracture/pipe joint.
Figure 6.12 Forces resulted from bending movement of the host-pipe (Allouche and Alam
2012)
Test specimens were prepared using 150 mm diameter, 1.2 m long cast iron host pipes, which
were in service for 70 years. The cast iron pipe was cut into two halves, which were connected
together with a thin circular wooden spacer placed in between. Then the samples were lined
with Aqua-Pipe liner, with a mechanical clamp holding the two sections in place.
To prevent rotation of the capped test specimen for the bending test, two custom-built supports
were designed and manufactured (Figure 6.13). A MTS servo-controlled actuator was
employed to impose a concentrated force to the ring fracture. A digital spirit level attached at
the crown of the host-pipe was used to measure the angular displacement.
34
Figure 6.13 Host pipe placed on custom built support for bending test (Allouche and Alam
2012)
The liner was found to be able to maintain its structural integrity even after the host pipe failed.
In one test, the deformed liner with a vertical deflection of 127 mm and an internal pressure of
0.8 MPa was left pressurised for one hour and there were no visible signs of leakage. However,
when there was no internal pressure in the lined pipe, the liner buckled at the invert during the
bending test.
To determine whether the liner will survive the host pipe failure, Stetter et al. (2017) conducted
three point bending tests on the pressurized lined pipes (Figure 6.14). The tests were conducted
on the exhumed AC pipes lined with Aqua-Pipe and Starline STRUCTURE-W liners. In the
tests, the CIPP liners were tested under internal pressures of both 400 kPa and 1000 kPa.
Figure 6.14 CIPP break test apparatus (Stetter et al. 2017)
After testing, it was found that all 10 lined pipe samples survived without liner failure. It was
observed that Aqua-Pipe and Starline STRUCTURE-W liners de-bonded from the host pipe
when the host pipe failed, but were able to maintain the internal water pressure. It was
concluded that each liner type is able to survive a circumferential failure without leaking in a
degraded AC pipe.
6.1.2.3 Pipe shear tests
Allouche and Alam (2012) investigated the behaviour of Aqua-Pipe Liner subjected to shearing
(Figure 6.15) under both pressurised and non-pressurised conditions at the location of ring
fracture/pipe joint. The test specimens were the same as those for bending tests (6.1.2.2).
35
Figure 6.15 Forces resulted from shear movement of the host-pipe (Allouche and Alam 2012)
Half of the test specimen was fixed while the other half was allowed to move only normal to
the pipe longitudinal axis (Figure 6.16). The fixed half was positioned inside two steel C-
channels to ensure its fixed boundary condition while the other half was placed inside a steel
box, bolted onto four guiderails as shown in Figure 6.18.
Figure 6.16 Experimental setup of the shear test (Allouche and Alam, 2012)
For the pressurised condition, the pipe was pressurized to an internal pressure of 0.4 MPa and
the actuator moved at an increment of 6.35 mm. It was found that when the displacement
reached 23 mm, the host pipe started to break, which resulted in a drop of the pressure. When
the displacement reached 38 mm, the host pipe started to crack in the pipe cross section and
de-bonding occurred between the liner and the host pipe. After the host pipe cracked, the liner
was found to be able to withstand the 0.4 MPa internal pressure. The liner eventually burst at
a lateral displacement of 89 mm and an internal pressure of 0.69 MPa. For the non-pressurised
condition, the actuator was stopped at an increment of 6.35 mm for inspection of the host pipe
and liner. During the testing, it was found that the liner at the ring fracture de-bonded from the
host pipe at the spring line when the vertical displacement reached 50 mm. The gap between
the internal surface of the host pipe and the external surface of the liner increased with the
36
increase of the displacement and the gap grew from the spring line region to the invert. When
the displacement reached 89 mm, a complete de-bonding was observed between the liner and
the host pipe at both the spring line and the invert. During a shear test under no internal pressure
sideways deformation was visible, both scenarios resulting in a reduction in the cross-sectional
area at the location of host-pipe ring fracture.
6.1.2.4 Pipe tensile tests
Allouche and Alam (2012) investigated the behaviour of Aqua-Pipe Liner subjected to tension
(Figure 6.17) under non-pressurised condition. The test specimen was the same as those for
bending tests.
Figure 6.17 Forces resulted from tensile movement of host-pipe (Allouche and Alam 2012)
One end of the capped specimen was fixed to a frame to simulate a fixed end, while the other
end was connected to a servo-control hydraulic actuator (Figure 6.18). Two parabolic-shaped
cut rods were welded to each cap. The actuator was pulled away from the fixed end at a rate of
3 mm/min until the host pipe was observed to slip out of the liner at the ring fracture.
Figure 6.18 Specimen restrained at one end (left) and pulled at other end (right) (Allouche
and Alam 2012)
37
The outer diameter of the liner is 176 mm and the contact length between the liner and the host
pipe is 1.2m. Experimental results showed that a force of 51 kN was required to overcome the
friction between the liner and the host pipe. This is corresponding to a friction value of
approximately 0.09 MPa (note that the contact area is 0.57 m2). This indicates a high degree of
friction at the interface between the liner and the host pipe in the tensile test. This might be due
to the mechanical interlock caused by the resin filling the corrosion pits in the internal surface
of the host pipe. It is noted that no change in the cross-sectional area was observed during the
tensile test.
Argyrou et al. (2017) performed a series of axial tension tests (Figure 6.19) to study the pull-
out capacity and investigate the failure mechanisms of pipelines with circumferential cracks or
leaking joints. The pipelines were lined with Starline2000® liners. In the tests, the host pipe
was made of ductile iron (DI) with 175 mm outer diameter 7.6 mm wall thickness. There was
a 3.3 mm thick interior cement mortar lining installed in the host pipe. The specimens for the
axial tension tests was made by either two straight DI pipe sections separated by a 6 to 12 mm
gap or two sections connected with a bell-and-spigot joint. One end of the pipe was clamped
to the test frame to simulate the fixed end, while the other was connected to a hydraulic
actuator.
Figure 6.19 Experimental setup for axial pull tests (Argyrou et al. 2017)
Test results summarised in Table 6.17 indicated that the behaviour of the lined pipes was
significantly affected by internal pressure. The presence of internal pressure reduced the de-
bonded lengths and hence, thus the axial deformations compared to the cases of no or very low
internal pressure. For the tests with no internal pressure, extensive de-bonding between the
liner and the host pipe occurred but there was no liner rupture.
38
Table 6.17 Axial tension tests results (Argyrou et al. 2017) Specimen
No.
Specimen
Type
Length at
each side
(m)
Loading
Rate
(mm/mm)
Pressure
(kPa)
Max
Force
(kN)
Opening
at test
end (mm)
Lining
Rupture
Debonding
Length
(mm)
G1 Gap 1.52,1.52 5.1 0 47.3 280 No 2845
G2 Gap 1.52,1.52 5.1 0 51.8 182 No 1880
G3 Gap 1.52,1.52 5.1 517 81.2 45 Yes 381
G4 Gap 1.52,1.52 5.1 517 88.2 78 Yes 559
J1 Joint 1.83,2.74 510* 517 92.1 71 Yes 508
J4 Joint 1.83,2.74 1.3 310 58.2 117 Yes 673
J5 Joint 1.83,2.74 2.6 310 89.4 100 No 1011
* This value is too large and may be an error in the original paper.
6.2 Long-term tests
Thermosetting resins or polymers exhibit reduction in strength over time similar to that
experienced by PVC pipes. Therefore long-term testing is crucial to determine long-term
properties of CIPP. The long-term study of pipe liners can be classified into long-term pipe
pressure testing, standard creep testing and accelerated creep testing.
6.2.1 Long-term pipe pressure testing
Straughan et al. (1995) conducted research on the long-term structural behaviour of cured-in-
place pipe (CIPP) and fold-and-formed pipe (FFP) liners made by different manufacturers
under external hydrostatic pressure similar to the field condition in a partially deteriorated
sewer. Each test specimen remained under a constant pressure for up to 10,000 hours or until
failure (similar to a hydrostatic design basis test). Test results indicated that creep leads to
buckling of the liners under significantly lower pressures than the initial buckling pressure. It
was also found from the regression analysis of the experimental results that the predicted long-
term buckling pressure is generally greater than that predicted by ASTM F1216 (1993).
Barbero and Rangarajan (2005) proposed a testing methodology to investigate the creep
behaviour of encased polymer and felt-reinforced polymer liners used in rehabilitating sewers.
Long-term tests were carried out on liners installed in steel pipes. Three specimens for each of
five liners (a total of 15 specimens) were tested under a constant external pressure for 19,000
hours. During the testing, a thermocouple was used to monitor the temperature of the liners.
39
Strain data were collected from a data acquisition system and were compensated for differences
in temperature, initial deformation, and coefficient of thermal expansion. Several viscoelastic
models were investigated in order to fit the data. The viscoelastic model were found to fit the
data well and was used to predict the long-term modulus used in the design of CIPP liner for
sewer rehabilitation.
Allouche and Moore (2005) undertook a long-term internal pressure test under steady
operating conditions. The specimen was a bell and spigot segment with a fire hydrant
connection, cut from a 152 mm diameter cast iron water main lined with the Aqua-Pipe liner
in 2003. The test partially followed ASTM D1598 (2002). An internal pressure of 1.2 MPa for
was planned to be applied for a period of 1000 hours. However after 450 hours the test was
stopped due to water leakage. An examination of the collected data revealed that the
displacement of the liner under 1.2 MPa internal pressure on Day 19 was nearly identical to
that recorded on Day 1. Therefore, no creep deformation was identified for this test.
Guan et al. (2007) investigated the effect of cyclic loading on the liner. Four PVC specimens
with an internal diameter of 150 mm were lined with an Aqua-Pipe liner. Several steel rings
were machined to a diameter greater than the PVC pipe’s spigot and a circular opening was
machined in the steel ring to expose the opening in the host pipe. The test set-up was shown in
Figure 6.20. The internal pressure was controlled to be between 0.4 and 0.8 MPa. The loading
frequency used in the testing was 40 s with 20 s at 0.4 MPa and 20 s at 0.8 MPa. More than
9000 cycles were conducted for the test, which was equivalent to 8.2 years with 3 surge events
per day. It was found that cyclic loads contributed to the liner’s displacement and plastic strain
increase in a liner under stresses well below that produced by the short-term burst pressure and
within the normal operation range for water pipes. It was also found that the effect of cyclic
loading can be accounted for by calculating the secondary creep increase based on the expected
maximum surge pressure.
The Trenchless Technology Center (2013) at Louisana Tech University conducted leak test of
sealed service connections in pressure pipes lined with CIPP liner. The testing included 5
specimens of 150 mm diameter PVC pipe lined with Aqua-Pipe liner. Each pipe specimen was
cut into three pieces and assembled into a continuous pipe with mechanical clamps. For each
pipe specimen, a lateral service connections was installed in the middle pipe piece while end
caps were installed to the two side pieces. The internal pressure was increased with two
increments of 0.34 MPa and one increment of 0.17 MPa until the target pressure of 0.86 MPa.
40
The test was run for 1000 hours (40 days). Observations showed no leakage in the location of
service connections in any of the specimens.
Figure 6.20 Experimental setup for cyclic testing
Figure 6.21 Glass reinforced CIPP long-term hydrostatic strength (Adapted from Microbac
(2009))
Microbac (2009) carried out hydrostatic design basis (HDB) testing on an Insituform liner
product InsituMainTM, following ASTM D2992 (2012). This test involved the hydrostatic
30
35
40
45
50
55
0.01 0.1 1 10 100 1000 10000 100000
Stre
ss (M
Pa)
Time (Hours)
Failure points
Predicted
Lower 95% Confidence Interval
Lower 95% Prediction Interval
41
pressure testing to failure of a minimum number of 18 pipe specimens at a variety of constant
(static) pressure levels over a period of 10,000 hours. The pipe specimens with the lowest glass
loading expected to be used in pressure pipe application were selected for testing. The HDB
testing results (Figure 6.21) indicated that a 50 year allowable tensile stress of 31.5 MPa, which
represents a 70% retention of initial tensile properties (45 MPa).
6.2.2 Standard creep testing
Standard creep testing were conducted by researchers with controlled temperature and
humidity, to investigate the long-term behaviour of the CIPP liner. All the creep tests in
literature discussed here followed ASTM D2990 (2001). Most of previous research focused on
the determination of creep retention factor, defined as the ratio of 50 years predicted modulus
to short-term modulus.
Lin (1995) and Straughan et al. (1998) conducted tensile, compression, and flexural creep tests
performed on flat samples (8 tensile specimens, 8 compressive specimens and 8 flexural
specimens) using sewer cured-in-place pipe (CIPP) liners for 3,000 hours. Another 4 curved
compressive specimens were tested to investigate the curvature on the creep behaviour. All
tension, compression, and bending tests were conducted in an environmental chamber with
controlled temperature and humidity. The recorded temperature ranged from 20 to 25˚C, and
the recorded humidity was from 55 to 75%. Following ASTM D2990, four stress levels were
considered for each type of the specimens. For bending tests, the stress levels were 6.89 MPa,
13.79 MPa, 20.68 MPa and 27.58 MPa. For tensile tests, the stress levels were 6.89 MPa, 10.34
MPa, 13.79 MPa and 17.24 MPa. For compressive tests, the stress levels were 6.89 MPa, 13.79
MPa, 20.68 MPa and 27.58 MPa. The creep strain was obtained as a function of time for each
specimen and Findley’s equation was employed for data fitting to predict the behaviour of the
CIPP liner. No significant difference was found between the axial strains of the flat and curved
specimens under compression. In addition, the average creep modulus at 3,000 hrs was found
to be 44%, 38% and 65% of the short-term elastic modulus for flexural, tensile and compressive
creep tests respectively.
Knight and Sarrami (2006) and Knight et al. (2018) performed flexural creep tests on a
reinforcement fabric impregnated liner with epoxy and/or vinyl ester resin for 10,000 hours.
Six test specimens were cut from flat plates. The applied initial stress the test was 25% of yield
stress. The creep rate, defined as the slope of the creep modulus versus the log time, was
42
observed to be linear in the first 100 hours. After this, the creep rate became constant. Results
showed that the liner has a short-term modulus of 1663 MPa and a 50-year modulus of
approximately 301 MPa, indicating a creep retention factor of 0.18.
Guan et al. (2007) conducted tensile creep tests on Aqua-Pipe CIPP liners. Ten dog-bone shaped
specimens were cut and tested under various stress levels for 5000 hours based on ASTM D638
and ASTM D2990. Five stress levels were considered: 2 MPa, 5.1 MPa, 8.2 MPa, 10.8 MPa, 14
MPa, which are 8.3%, 21.7%, 34.8%, 45.7% and 59.5% of yield stress respectively. Two
specimens were considered for each stress level. All the tests were conducted in an environmental
chamber with a temperature of 21.1 ± 0.5 ˚C). The creep strain as a function of time for each
specimen under a given stress level were fitted to three strain creep models, namely the Power Law
creep, Exponential Law creep and the Eight-Parameter creep model. The Eight-Parameter creep
model was found to be the best representation of the experimental creep data.
Microbac (2011) performed flexural creep tests on specimens cut from a flat epoxy plate with a
single glass layer following ASTM D2990 (2001). Five specimens were selected and tested. The
applied stress was 0.25% of the short-term flexural modulus. The tests were performed at 23 ± 2°C
temperature and 50 ± 5% relative humidity over a period of 10,000 hours. The long-term modulus
was extrapolated based on the most linear portion of the data. It was found that the 50 year creep
retention factor was 0.54 while the 100 year one was 0.51.
Riahi (2015) conducted flexural creep tests on the CIPP liners with Alpha Owens-Corning
(AOC) and Interplastic Corporation resins. Five specimens were tested for each resin type. The
tests lasted for 96,000 hours equivalent to 11 years. Results showed that the 50-year creep
retention factor extrapolated based on the 10,000 hours of data is in general higher than that
based on the 96,000 hours of data and the creep retention factor of 0.5 used in industry may
not be conservative.
6.2.3 Accelerated creep testing
Estimating the long-term deformation under static load would require conventional creep
experiments over a long period of 5–10 years in order to account for product lifetimes of 50–
100 years.
As an alternative to extrapolation, accelerated creep testing can be carried out at low
stress/temperature levels, in such a way that the long-term creep and creep-rupture properties
43
can be determined within shorter time scales. The creep rate is accelerated, thus reducing the
time needed for a given amount of creep to occur. As a result, failure of the specimen can be
achieved in practical timescales.
The fundamental basis for accelerated creep testing is that a material’s resistance to creep can
be overcome by providing energy, which can be in the form of heat or stress. Most of the
methods are based on time-temperature-stress superposition principle, or some variation, which
involves the manipulation of temperature, applied stress, or both as a way to reduce the testing
time period.
Time-temperature superposition principle (TTSP)
It has been recognised that time and temperature are equivalent in the way they affect creep
properties (Ferry 1980). By relating these effects, it is possible to predict long-term properties
of polymers from short-term tests carried out at temperatures higher than those encountered in
the field conditions. This is the fundamental basis for the time-temperature superposition
principle (TTSP) (Leaderman 1943; Tobolsky and Andrews 1945; Seitz and Balazs 1968). It
should be noted that this method was developed for predicting linear viscoelastic properties of
homogeneous polymers and may not be accurate when applied to semi-crystalline polymers
(Lai and Bakker 1995).
In this method, multiple specimens are tested under a constant load at different temperatures
resulting in separate plots of creep strain versus log time at different temperatures. A reference
temperature is then chosen, usually close to the ambient temperature, and all individual curves
are shifted along the log (time) axis to compensate for different temperatures. By applying the
principle of superposition a creep master curve can be generated.
Farrag and Shirazi (1997) and Farrag (1998) determined creep properties of high-density
polyethylene geogrid using the TTSP with 1000 hours at different temperatures. The master
curve generated from the TTSP data were compared with the creep curve obtained from the
conventional test method and a good agreement was found.
Stepped isothermal method (SIM)
The stepped isothermal method (SIM) (Thornton et al. 1998) can be considered as a special
case of the TTSP method. It is a short-term creep test in which the temperature is increased in
44
steps and the mast curve can be produced based on a single experiment. After testing, the
measured strain is first re-scaled and then shifted according to the time–temperature
superposition principle to create a master curve. The SIM master curves have been found to
match those from the classical TTSP procedure while reducing the experimental effort to a
minimum.
In a SIM test, a single specimen is loaded under a constant load. The temperature is then
increased in steps, either until sufficient creep is achieved or the specimen fails. A series of
corrections are required to account for the temperature stepping and different creep degrees at
each stage, before the master curve is obtained to predict the long-term behaviour. It should be
noted that the SIM tests rely on an activation energy which needs to be constant at all
temperature levels to ensure the same creep mechanism.
The SIM was mainly applied for testing geosynthetics. Several investigations for various
geosynthetics proved a good agreement of SIM master curves with conventional long-term
creep tests (Zornberg et al. 2004; Bueno et al. 2005; Yeo and Hsuan 2010). Therefore, the SIM
became a well-established method for the accelerated product testing of geosynthetics and
standards were published for both tensile creep (ASTM D6992 2016) and compressive creep
(ASTM D7361 2018).
Despite the advantages of the SIM, only a few attempts have been made to extend the
application of this method to other polymers. Alwis and Burgoyne (2008) applied the SIM
successfully to creep testing of aramid yarns. Similar to geosynthetics, yarns exhibit a very
small cross section and can therefore be heated quickly. Recently, attempts were made for
polymers with large thickness. Thomas et al. (2010) investigated High Density Polyethylene
(HDPE) under tensile loading for pressure pipes. Bozorg-Haddad and Iskander (2011) studies
the HDPE subjected to compressive loading. Achereiner et al. (2013) successfully applied it to
investigate the creep behaviour of polypropylene up to approximately 100 years.
Time-stress superposition principle (TSSP)
Accelerated creep testing can also be achieved by supplying the energy with increased stresses.
In the time-stress superposition principle (TSSP) (Lai and Bakker 1995), multiple specimens
are required to be tested at ambient temperature at various stress levels. Each stress level will
result in a separate plot of creep strain versus log (time). After a reference stress is chosen,
45
similar to the TTSP, all individual curves are shifted along the log time axis to compensate for
different stresses. Then a master curve is produced based on the principle of superposition.
This method was used by researchers to predict the creep behaviour of various polymers. Lai
and Bakker (1995) applied the TSSP on HDPE at various stress levels and ambient temperature
and derived a unified creep relation taking into account the physical aging effect. Hadid et al.
(2004) performed flexural creep testing on fibre glass reinforced polyamide materials at
different stress levels, in which each test lasted for 30 minutes. The creep master curve was
created based on an improved empirical model in which non-linear behaviour was accounted
for. Jazouli et al. (2005) conducted non-linear creep tests on polycarbonate to investigate the
stress induced changes in intrinsic timescale at room temperature. Based on the concept of
time–stress equivalence, the creep compliances were determined as a function of time at nine
different stress levels and shifted along the logarithmic time axis to obtain a master compliance
curve at a given reference stress level. It was found that the TSSP provides an accelerated test
technique for evaluating the material’s long-term mechanical properties. Starkova et al. (2007)
studied the tensile creep behaviour of polyamide 66 and its nanocomposites filled with 1% by
volume TiO2 nanoparticles that were 21 nm and 300 nm in diameter. A master curve was
created and employed to predict the long term behaviour of the tested material.
Stepped isostress method (SSM)
Giannopoulos and Burgoyne (2011) proposed a new accelerated technique, termed the stepped
isostress method. This method is similar to the stepped isothermal method, but the acceleration
is achieved by increasing the stress rather than the temperature. Compared with SIM, this
method is advantageous as elevated temperatures are not required, which may affect the
material chemical properties. In addition, the problem concerning the slow and non-uniform
heating of thick samples is avoided.
In this method, a single specimen, in contrast to the many specimens required by the TSSP, is
loaded under various stress levels. A step-wise increase in stress was applied to this single
specimen at a constant temperature. At each stress step, a creep strain versus time curve is
obtained and then adjusted to compensate for different stress levels. After that, a creep master
curve at a reference stress level is created for predictions.
46
This method was applied by Giannopoulos and Burgoyne (2011) to predict the long-term
behaviour of Kevlar 49 yarns. The creep curves and rupture times from the SSM were
compared with those from the SIM and conventional creep tests from literature. Good
agreement among the results from the three test methods were found. Giannopoulos and
Burgoyne (2012) conducted accelerated creep testing on high modulus aramid fibres. The tests
showed good agreement between the SIM and the SSM tests for Kevlar 49 above 60% of the
average breaking load and Technora. However, a significant difference was found in the creep
rupture life predicted for Technora, especially at low stress levels. Hadid et al. (2014) employed
the SSM to investigate its application to thick thermoplastic specimens. Excellent match was
observed between the master curves produced by the classical TSSP method and those by the
SSM.
6.2.4 Mechanical aging
Stewart et al. (2015) performed full-scale tests to examine the performance of cast iron (CI)
gas pipelines lined with cured-in-place liners (Starline® 2000 PSE-35 liner), which had been in
service for 10 to 16 years. The test specimens were two sections of lined 150 mm diameter CI
pipe and two sections of lined 300 mm diameter CI pipe with joints. All these four lined pipe
sections were about 2.4 m long with a joint in the middle. The joints were used to simulate
either circumferential cracks or weak and degraded joints in the field. Two sets of mechanical
tests including flexural and axial compression/tension test were performed on the CI
specimens.
Four-point flexural tests were conducted to simulate traffic loading over two 50-year service
life cycles (a total of 100 years). The displacements and joint rotations considered in the
flexural tests were consistent with those determined from the analytical models developed at
Cornell, which were validated by full-scale field tests (assuming typical soil and flexible
pavement conditions in the field). Undermining and backfill also were simulated in the tests.
Axial compression/tension tests were also carried out to simulate the impact of yearly
temperature changes on the lined pipe over the two 50-year service life cycles (a total of 100
years). Operating pressures of 0.102 MPa and 0.414 MPa were applied in the pipe specimens
with diameters of 150 and 300 mm respectively.
The testing consisted of the following stages: traffic loads/bending cycles (1 million cycles);
Undermining excavation event; Additional traffic loads/bending cycles (100,000 cycles);
47
Thermal expansion/contraction cycles (50 cycles); Traffic loads/bending cycles (another 1
million cycles); Excavation event; Additional traffic loads/bending cycles (another 100,000
cycles); Thermal expansion/contraction cycles (another 50 cycles); Post-testing verification
pressure.
It was observed that none of the specimens leaked during all stages of the mechanical aging
tests. After the mechanical aging tests, the internal pressure of a 150 mm diameter pipe
specimen was increased to 1.034 MPa and the lined pipe continued to maintain pressure
tightness. For the 300 mm diameter specimens, holes were drilled through the bell and into the
gap between the spigot end and the liner. The each specimen was pressurised to 0.620 MPa
and the liner did not leak. After all the testing, the pipe joints were cut in the longitudinal
direction for inspection. For the 150 mm diameter specimens, De-bonding at the separation
between bell and spigot occurred at the joints of. For the 300 mm diameter specimens, some
minor liner damage were observed, but no leakage was found.
6.3 Summary 6.3.1 Short-term testing
For standard specimen testing, most of the previous research focused on the material properties
of CIPP liners installed in sewer pipelines, in particular the tensile and flexural properties of
liners. The results are summarised in Table 6.18. A limited number of studies were conducted
to determine the material properties of CIPP liners for pressure pipes, which are presented in
Table 6.19. For liners in pressure pipes, attention was mostly paid to tensile properties, as
tension will be mostly dominant in this case. It should be noted that for pressure pipes, flexural
properties should also be determined as the liner may experience negative pressure conditions
or external pressure, or be subjected to bending for small diameter pipelines, in which cases
flexure will be dominant for the liner.
Some other researchers also carried out experiments to investigate the bond strength between
the liner and the host pipe, e.g., the lap shear test, peel test (for Starline liners) and adhesion
test (for Aqua-Pipe and Starline liners). Split disk tests were also conducted to examine the
effect of wavy imperfections on the strength of a liner (Ampiah et al. 2008; 2010).
For large-scale pipe tests, most of the research found from literature was on Aqua-Pipe. It was
found that Aqua-Pipe liners passed pressure tests, negative pressure tests, bending, shear, and
48
tensile tests (Allouche et al. 2005; Allouche and Moore 2005; Matthew et al. 2012; Stetter et
al. 2017). However, very limited studies were on other CIPP liners. Only tensile testing was
found for Starline2000® liners and bending test for Starline STRUCTURE-W liners. Testing
results for large-scale pipe testing can be seen in Table 6.20. Therefore, further testing on other
CIPP liners will need to be conducted.
49
Table 6.18 Material properties of CIPP liners for old sewer pipes
Reference
Host Pipe
/Diameter
(mm)/
Location
Liner
Age
(years)
Liner
Tensile
Strength
(MPa)
Tensile
Modulus
(MPa)
Tensile
elongation
(%) at
break
Flexural
Strength
(MPa)
Flexural
Modulus
(MPa)
Durometer
(Shore D)
Hardness
Test
(ASTM
D2240)
Water
Tightness
Buckling
pressure
(kPa)
Specific
Gravity
(g/cm3)
Ovality (%)
Lystbaek
(2007)
14
39.2 ± 2.5 2603 ± 140
5 41.0 ± 5.6 3308 ± 612
0 41.5 ± 4.7 3480 ± 266
Interplastic
Corporation.
(2008)
- 0
Resin/felt
composites
(Field)
21.93/1.02 4482/225 - 45.6 ±
3.97 3796 ± 139
- 0
Resin/felt
composites
(Lab)
26.9/2.52 4550/89 - 62.9/9.42 4240/631
Allouche et
al. (2012) &
Allouche et
al. (2014)
Clay pipe
/203/
Denver
25 Insituform 21 ± 1.2 2838 ± 280
1.5-4.5 47 ± 3.8 2313 ± 125 Inner 58.9
Outer 77.0 1.16 7.4
16 ± 1.4 - 48 ± 3.1 3406 ± 295
Brick sewer
/1219/
Denver
23 Insituform
22 ± 1.5 2943 ± 403
2.5-9.5
34.7 ± 4.5 1259 ± 159
Inner 46.6
Outer 62.7 1.07 -
42 ± 6.1 1818 ± 485
21 ± 1.6 2636 ± 415 48.5 ± 2.4 2089 ± 168
8 16 ± 1.2 1.5-9.0 48 ± 2.8 3378 ± 276
Clay pipe
/203/
Columbus
5 Reynolds
Inliner
27 ± 2.9 2500 ± 301 1.0-11.0 44 ± 14.1 2386 ± 343 Inner 62.7
Outer 81.4 1.11 5.07
0 - - 50 ± 3.4 -
Brick sewer
/914/
Columbus
21 Insituform 20 ± 1.7 2174 ± 293 2.5-6.0 42 ± 2.7 1426 ± 200 Inner 64.8
Outer 78.6 1.17 -
Matthew
(2014) 0
UV cured CIPP
liner (Reline 147 ± 25
14000 ±
4400 200 ± 66
10200 ±
3900 Passed 455 1.46 ± 0.04
50
Reference
Host Pipe
/Diameter
(mm)/
Location
Liner
Age
(years)
Liner
Tensile
Strength
(MPa)
Tensile
Modulus
(MPa)
Tensile
elongation
(%) at
break
Flexural
Strength
(MPa)
Flexural
Modulus
(MPa)
Durometer
(Shore D)
Hardness
Test
(ASTM
D2240)
Water
Tightness
Buckling
pressure
(kPa)
Specific
Gravity
(g/cm3)
Ovality (%)
Vitrified
clay pipe
/250/ Frisco
America Blue-
Tek™ liner)-
longitudinal Inner 71.0
± 1.4 Outer
64.2 ± 3.7
UV cured CIPP
liner (Reline
America Blue-
Tek™ liner)-
circumferential
390 ± 93 13100 ±
3400
Reinforced
concrete
pipe /2400/
Irving
0
Water-Cured
CIPP liner
(Insituform
iPlus®
Composite
Liner)
80 ± 9 6900 ± 300
Inner 50.3
± 3.9 Outer
67.9 ± 0.4
1.258 ±
0.0015
Sterling et al.
(2016)
/250/
Edmonton 19 22 3011 42 2285
Inner 68.6
Outer 78.9 83 1.25 2.7-4.3
/200/
Edmonton 19 25 3517 47 2515
Inner 68.2
Outer 79.2 138 1.25 4.5-5.75
/525/
Houston 17 24 3208 48 2328
Inner 61.2
Outer 61.3 - 1.17 1.4
/450/
Houston 17 22 3110 50 2334
Inner 65.4
Outer 75.7 - 1.18 1.7
/1050/
Indianapolis 25 19 2422 32 1636
Inner 57.0
Outer 65.7 - 1.08 -
51
Reference
Host Pipe
/Diameter
(mm)/
Location
Liner
Age
(years)
Liner
Tensile
Strength
(MPa)
Tensile
Modulus
(MPa)
Tensile
elongation
(%) at
break
Flexural
Strength
(MPa)
Flexural
Modulus
(MPa)
Durometer
(Shore D)
Hardness
Test
(ASTM
D2240)
Water
Tightness
Buckling
pressure
(kPa)
Specific
Gravity
(g/cm3)
Ovality (%)
/200/
Nashville
Dunston
19 24 2591 47 2080 Inner 65.2
Outer 72.2 - 1.14 3.7
/200/
Nashville
Wyoming
9 18 2764 38 1948 Inner 64.6
Outer 67.4 - 1.21 3.6
/375/ NYC 23 26 3821 55 3293 Inner 73.3
Outer 72.1 - 1.31 -
/300/ NYC 24 23 2237 50 1966 Inner 57.7
Outer 58.7 - 1.15 -
/300/
Northbrook 34 30 2989 54 2223
Inner 65.6
Outer 76 34 1.19 0.33-0.38
/750/
Winnipeg
Richard
34 - 59 3117 Inner 57.4
Outer 65.8 - 1.21 -
/450/
Winnipeg
Kingsway
34 - 47 2233 Inner 54.1
Outer 60.9 - 1.14 -
/750/
Winnipeg
Mission
28 - 31 1694 Inner 57.3
Outer 64.9 - 1.07 -
Note: Empty cells in the table means that there is no information available.
52
Table 6.19 Material properties of CIPP liners for pressure pipes
Reference Host Pipe
(Diameter) Liner Age Liner
Tensile
Strength
(MPa)
Tensile
Modulus
(MPa)
Yield
Strain
%
Flexural
Strength
(MPa)
Flexural
Modulus
(MPa)
Durometer
(Shore D)
Hardness Test
Barcol
Hardness
Test
Specific
Gravity
(g/cm3)
Ovality
(%)
Poisson’s
ratio
Allouche et
al. (2005) Cast iron
(150 mm) 0 Aqua-Pipe 55
2000 MPa
(subsequent
modulus 180)
1.3
Brown et al.
(2008)
Cast iron
(150 mm)
0
Composite liner
(Exhumed-
longitudinal)
61 ± 0.6 2019 ± 8.6 1
0
Composite liner
(Fabricated-
longitudinal)
61.3 ± 2.8 2017 ± 243 0.9
0
Composite liner
(Fabricated-
circumferential)
88.4 ± 4.7 3040 ± 120 0.9
Matthew et a.
(2012c) Cast iron
(150 mm) 0
Aqua-Pipe
(Exhumed-
longitudinal)
65 ± 2.1 3559 ± 1054 55 ± 4 2530 ± 105
Inner surface
39.8 ± 5.4 Outer
surface 60.8 ±
1.5 (ASTM
D2240)
Inner surface
1.9 ± 0.3
Outer surface
9.4 ± 1.4
1.154 ±
0.93 2.5
Stewart et al.
(2015)
Cast iron
(150 mm)
Field aged (16
years) Starline®2000 PSE-
35 liner
(longitudinal)
131.2 ± 9.7
760
Field (16 years)
and Mechanically
(Equi. to 100
years) aged
118.8 ± 3.5
Field aged (16
years) Starline®2000 PSE-
35 liner
(circumferential)
24.8 ± 1.2
Field (16 years)
and Mechanically 23 ± 0.9
53
Reference Host Pipe
(Diameter) Liner Age Liner
Tensile
Strength
(MPa)
Tensile
Modulus
(MPa)
Yield
Strain
%
Flexural
Strength
(MPa)
Flexural
Modulus
(MPa)
Durometer
(Shore D)
Hardness Test
Barcol
Hardness
Test
Specific
Gravity
(g/cm3)
Ovality
(%)
Poisson’s
ratio
(Equi. to 100
years) aged
Cast iron
(300 mm)
Field aged (10
years) Starline®2000 PSE-
35 liner
(longitudinal)
80.5 ± 6.3
Field (10 years)
and Mechanically
(Equi. to 100
years) aged
79.6 ± 5.1
Field aged (10
years) Starline®2000 PSE-
35 liner
(circumferential)
40.8+2.4
Field (10 years)
and Mechanically
(Equi. to 100
years) aged
44.1 ± 2.6
Zhong (2015)
Ductile iron
(150 mm) 0
InsituMain
(longitudinal) 42.1 3316.4
InsituMain
(circumferential) 77.2 4874.6
Starline®2000
(longitudinal) 51.6±10.9 2599±0.5
Starline®2000
(circumferential) 0.01±0.003* 958.4±0.2
Yan (2016)
(305 mm) 0
A new composite
liner (longitudinal) 31.9±7.7 470.1±43.5 4 25
A new composite
liner
(circumferential)
25.4±10.3 178.7±70.3
54
Reference Host Pipe
(Diameter) Liner Age Liner
Tensile
Strength
(MPa)
Tensile
Modulus
(MPa)
Yield
Strain
%
Flexural
Strength
(MPa)
Flexural
Modulus
(MPa)
Durometer
(Shore D)
Hardness Test
Barcol
Hardness
Test
Specific
Gravity
(g/cm3)
Ovality
(%)
Poisson’s
ratio
A new composite
liner (45 degree
from pipe axis)
4.0±0.9 67.2±31.9
Microbac
(2011) InsituMain 76.2 ± 2.3
3178 ±
34.5
South East
Water and
ALS
Industrial Pty
Ltd. (2017)
(100 mm) 0 Aqua-Pipe 74.3 ± 3.0
(AS 1145) 2586 ± 644 47.4 ± 1.5
(150 mm) 0 Aqua-Pipe 79.6 ± 2.4 2507 ± 174 51.6 ± 2.7 0.17
Note: Empty cells in the table means that there is no information available.
* The value is too small. There might be an error in the original literature (Zhong 2015).
55
Table 6.20 Testing results of pipe testing Reference Host pipe Liner Pressure test Negative pressure test Bending test Shear test Tensile test
Allouche et al. (2005)
and Allouche and
Moore (2005)
150 mm cast iron Aqua-Pipe Resisted pressure up to 3.8
MPa
Matthew et al. (2012) 150 mm cast iron Aqua-Pipe Withstood -0.1 MPa for
70 hours with very little
deflection
Allouche and Alam
(2012)
150 mm cast iron Aqua-Pipe Subjected to internal
pressure, was able to
maintain its structural
integrity even after the
host-pipe failed
Perform adequately, after
undergoing lateral
deformation equal to
50% of the pipe’s
external diameter at the
location of a ring fracture
A high degree of
friction at the liner-
host pipe interface
during uni-axial test
Argyrou et al. (2017) 175 mm ductile iron Starline2000®
liners
Response of the
lined pipes is
strongly affected by
internal pressure
Stetter et al. (2017) 100mm and 150mm Aqua-Pipe and
Starline
STRUCTURE-W
All 10 samples ‘passed’
Sanivar AG (2000) 150 mm and 300 mm Saniline W liners Resisted pressure up to 4.8
MPa for three holes of DN 32
mm in host pipe. Resisted
pressure up to 2.85 MPa for at
a joint. Resisted pressure up
to 2.6 for one hole of DN 100
mm in host pipe
Note: Empty cells in the table means that there is no information available.
56
6.3.2 Long-term testing
Based on the above literature review, it has been found that two long-term pipe pressure tests
were conducted for gravity sewer pipes, in which case external pressure was applied to the
pipes to determine the buckling pressure (Straughan et al. 1995; Barbero and Rangarajan 2005).
For pressure testing of internally pressurised water pipes, hydrostatic design basis (HDB)
testing was conducted by Microbac (2009) on InsituMainTM. A long-term pressure test under
steady operating conditions was undertaken by Allouche and Moore (2005). However, this test
was terminated after 19 days due to water leaking. The Trenchless Technology Center (2013)
also conducted leak test on Aqua-Pipe liner lined 150 mm diameter PVC pipe with service
connections subjected a maximum pressure of 0.86 MPa for 40 days and no leakage in the
location of service connections was found.
For long-term pressure testing of CIPP liners, it should be noted that the internal pressure for
water pipes is much higher than the external pressure for gravity pipes, which may make the
long-term pressure testing of water pipes impractical or difficult to conduct. In addition, it is
also expensive to carry out long-term pressure testing, even for gravity pipes with low external
pressure. For pressure pipes, the main difficulty is maintaining sustained high internal pressure
in the specimen for a prolonged time period. Instead, standard creep testing may be an
achievable alternative to investigate the long-term behaviour of the internally pressurised lined
pipes. However, the correlation between the long-term pressure testing and standard creep
testing should be established. According to ASTM F1216 (2016), for pressure pipes, the long-
term flexural and tensile strength are required for the design of partially and fully deteriorated
pipes, respectively.
For standard creep testing of pipe liners, most previous research concentrated on flexural creep
tests of CIPP liners considering different stress levels up to 96,000 hours. Limited number of
tests were conducted for tensile creep and only up to 5,000 hours were considered. As the creep
test results are extrapolated for 50 years, longer creep testing is required for better prediction.
In addition, previous research mainly focused on the determination of creep modulus reduction
over time and the creep strain change over time. Creep testing were performed considering
different stress levels up to 96,000 hours for flexural creep, 5,000 hours for tensile creep and
3,000 hours for compression creep and the results were extrapolated to predict the creep
modulus reduction at 50 years. It is recommended that further creep, creep rupture or HDB
57
tests be conducted on CIPP liners in further testing to determine the reduction in strength over
time.
However, from literature, it has been found that little has been known on the reduction of
tensile, flexural and bond strengths over time. As an alternative to standard creep testing,
different accelerated creep testing methods were developed and applied to different polymeric
materials, in order to reduce time for creep testing. Literature showed that there are in total four
methods, namely, the TTSP method, SIM method, TSSP method and SSM method. Among
these four methods, the SIM and SSM methods are most promising. The SIM method was
already applied to creep testing of material such as aramid yarns, HDPE, and polypropylene
while the SSM method was employed to conduct creep testing on the Kevlar 49 yarns,
Technora and thermoplastic materials. It is recommended that either SSM or SIM tests be
conducted on CIPP liners in further testing.
Mechanical ageing tests were performed by Stewart et al. (2015) to evaluate the performance
of the Starline CIPP lined cast iron pipes with a round crack. Both the effect of traffic loads
and temperature changes were taken into account. Results showed that after 10 (or 16) years of
field aging and 100 years of mechanical aging, the CIPP liner (Starline® 2000 PSE-35 liner)
still functions satisfactorily. Although there is de-bonding between the pipe and the liner or
some liner damage after aging, the liner still maintains its pressure integrity for a certain
pressure level. Based on this study, it could be concluded that mechanical aging may not
reduce the strength of the CIPP liners tested over the long-term.
6.4 Gap identification 6.4.1 Short-term testing
As discussed in section 6.3.1, both tensile and flexural properties are required for design and
analysis of CIPP liners for pressure pipes. From Table 6.19, it can be seen that tensile and
flexural properties were available for the liner products to be tested in the CRC-P project except
the Saniline liner. Quite large variations of the liner properties were observed from the testing
results. For example, the tensile strength of Aqua-Pipe liners could range from 55 MPa to 82.9
MPa while that of Starline liners could vary between 58.4 MPa and 134 MPa. The same applies
to flexural properties. In addition, the tensile strength values in longitudinal and circumferential
directions for different liner types showed different trends. The CIPP liner in Brown et al.
(2008) showed larger tensile strength in circumferential direction while that in Stewart et al.
58
(2015) exhibited larger tensile strength in longitudinal directions. In order to get better statistics
of the liner properties, more specimens in both longitudinal and circumferential directions still
need to be tested. In addition, all testings of material properties should be conducted on Saniline
liners, as no data were available.
For bond strength between the liner and host pipe, only limited adhesion/shear testing results
were available for Aqua-Pipe and Starline liners. In addition, parallel plate/ring tests were only
conducted on Aqua-Pipe liners and no results were found for Saniline, InsituMain or Starline.
Therefore, adhesion/shear tests may be required for InsituMain and Saniline liners while ring
tests may be needed for InsituMain, Saniline and Starline liners.
For large-scale pipe testing, all the required structural tests, including pressure tests, negative
pressure tests, shear, tensile and break tests were conducted for Aqua-Pipe. However, no
negative pressure tests, shear and tensile tests were found to be performed on Starline,
InsituMain and Saniline liners. In addition, no pressure tests have been found on Starline and
InsituMain liners and no bending tests were found on Saniline and InsituMain liners.
The gaps are briefly summarized in Table 6.21 as follows
Table 6.21 Gaps for short term testing
Short term tests Identified gaps
Tensile and flexural
properties
Due to large variations in material properties, more specimens in
longitudinal and circumferential directions still need to be tested to
get better statistics.
No experimental data of material properties is available Saniline
liners
Bond strength No results were found for Saniline and InsituMain.
Parallel plate/ring
tests
No results were found for Saniline, InsituMain or Starline.
Large-scale pipe
testing
No negative pressure tests, shear and tensile tests were found to be
performed on Starline, InsituMain and Saniline liners.
No bending tests were found on Saniline and InsituMain liners
59
6.4.2 Long-term testing
Due to the high cost and difficulty in maintaining sustained high internal pressure in the
specimens for a long period of time, it may be impractical/difficult to conduct long-term pipe
pressure testing. Instead, standard creep testing may be an achievable alternative to investigate
the long-term behaviour of the internally pressurised lined pipes. However, a correlation
between long-term pressure testing and standard creep testing needs to be established.
From literature, it can be seen that little study was conducted for tensile creep tests of CIPP
liners. Only up to 5,000 hours of creep was conducted and longer creep testing time is required
for better prediction of liner behaviours. According to ASTM F1216 (2016), for pressure pipes,
the long-term flexural and tensile strength are required for the design of partially and fully
deteriorated pipes, respectively, in addition to the long-term elasticity modulus. Therefore,
attention should be paid to the effect of creep on the strength (e.g. tensile and flexural) of
the liners. In addition, bonding between the liner and host pipe might be required at service
connections. As a result, long-term bond strength might also be needed.
As an alternative to standard creep testing, different accelerated creep testing methods were
developed and applied to different polymeric materials. However, no study has been
conducted to investigate the applicability of the SIM and SSM methods on CIPP liners
thus far.
7. Review of Numerical Modelling of corroded host pipes with
imperfect liners
Compared with experimental tests, numerical modelling is a relatively low cost alternative for
investigating the behaviours of lined degraded pipes. The numerical studies in the previous
research can be classified into the following five categories.
7.1 Effect of size and geometry of a defect on host pipe
To develop the relationship between the size and geometry of a hole in a host pipe and the burst
pressure of a liner, Guan et al. (2007) carried out numerical analyses of a host pipe with 230
mm external diameter and 12.7 mm thickness. Different geometries of the hole were
considered, including square, circle, rectangle and ellipse in both hoop and axial directions. All
60
the defect shapes were considered to have the same cross-sectional area. Results showed that
rectangle and ellipse defect shapes with the long axes in the pipe longitudinal direction show a
significantly lower burst pressure than a circular defect. However, as liner strength exhibits
anisotropy in different directions, the effect of the defect shape could change if the higher
strength direction fibre is aligned along the pipe longitudinal axis.
Brown et al. (2014) investigated the performance of a CIPP liner, which was installed in a hot
pipe with one section completed gone/missing. Three-dimensional orthotropic elastic analysis
was conducted on the liner with greater strength and stiffness in the hoop direction. Results
showed that due to the greater strength and stiffness in the hoop direction, the maximum axial
stresses in the liner were reduced by approximately 27% and the tensile stress in the liner was
governed by the unconfined hoop stress at the maximum operational pressure. The effect of
friction between the host pipe and the liner was also studied and it was found that the coefficient
of friction ranging from 0 to 0.577 has little effect on the numerical results.
Shou and Chen (2018) investigated the effect of a liner on the stress level in a host pipe. The
stress-strain behaviours of buried pipes with circular corrosion pits (non-through wall pits),
considering internal pressure and surface loads were simulated with the installation of a CIPP
liner. The friction coefficient between the host pipe and the liner was considered to be 0.46.
Results showed that a larger diameter to thickness ratio resulted in a greater increase of stress
and displacement at the corrosion pit. It was also found that the internal pressure mainly
controlled the stresses near the host pipe defect area, while the displacements were mainly
governed by surface loading. In addition, the stress and displacement were significantly
reduced at the damaged area by the installation of the CIPP liner.
7.2 Effect of material properties of the host pipe on the pressure rating of
liners
Guan et al. (2007) numerically simulated the bursting pressures of an internally lined host pipe
with a 230 mm external diameter and a 12.7 mm thickness. Two materials, namely, cast iron
and PVC, were considered for the host pipe and the circular defect size in the host pipe varied
from 12.7 mm to 203 mm. Results showed that the predicted liner burst pressures for cast iron
and PVC host pipes are quite different when the defect size is relatively small (e.g., the diameter
of the defect was smaller than 100 mm), but the burst pressure becomes similar as the defect
size grows larger. The explanation is that for small gap size the host pipe and liner work
61
together to hold the pressure and the host pipe restricts the deformation of the liner. When the
gap size becomes larger, the restriction of liner deformation by the host pipe decreases and
most of the pressure is taken by the liner in the gap area.
7.3 Effect of liner imperfections on the performance of the liners
Different liner imperfections have been studied by researchers to investigate their effect on the
performance of the liners.
Allouche et al. (2005) and Jaganathan et al. (2007) conducted comprehensive experimental and
numerical studies to quantify the effect of liner folds on the pressure rating of a CIPP liner. It
was found that when a longitudinal fold is transverse to a hole in the host pipe wall, the burst
pressure is the lowest. Through a parametric study using finite element analyses, Jaganathan et
al. (2007) established a relationship between the geometry of the fold and the critical burst
pressure for the liner, given a hole in the host pipe.
Zhao (2003) investigated the effect of the variation of liner thickness on the critical buckling
pressure of a liner. In this study, the thickness variations of the liner in both the longitudinal
and circumferential directions were assumed to be sinusoidal and the frequency and magnitude
of the thickness variations were studied. When the thickness only varies in the longitudinal
direction, it was found that the buckling pressure increases with the increase of the magnitude
of the thickness variation. When the thickness only varies in the circumferential direction, it
was found that the buckling pressure decreases with the increase of the magnitude of the
thickness variation. When the thickness varies in both directions, it was found that the buckling
pressure decreases with the increase of the frequency of thickness variation.
El-Sawy (2013) employed a finite element (FE) method to investigate the inelastic stability of
cylindrical liners with localised wavy imperfections under external pressures (grout pressure,
external water pressure or internal negative pressure). A smooth interface between the liner
and the host pipe was assumed. The liner was modelled using elastic-perfectly-plastic stress–
strain relationship to study the effects of the different geometrical parameters of the liner on its
stability. Results showed that an increase in the yield stress results in a decrease in the
imperfection size and liner thickness and convergence of the liner buckling to the elastic case
with normalised pressure close to unity. In addition, it was found that the liner’s normalised
critical pressure generally decreases with the increase of the imperfection angle.
62
In the oil and gas industry polymeric liners are used for corrosion resistance to metal pipes.
Gases derived from the oil can permeate through the liner wall, into the gap between the liner
and host-pipe. When rapid depressurisation occurs, the gases can generate an external pressure
on the liner, which may lead to liner buckling (Rueda et al. 2012). Rueda et al. (2012) simulated
the buckling collapse, including post-collapse pressure drop, of HDPE liners. In the analysis,
a very small out of roundness was introduced to initiate buckling. The contact between the steel
pipe and the HDPE liner was assumed to be frictionless and the pressure applied on the external
surface of the liner is considered to be uniform and linearly increasing. Three types of analyses,
namely, static with hydrostatic elements, Riks general static and general static, were
considered. Results showed that finite element models with hydrostatic elements reasonably
simulated the collapse of polymeric liners under external pressure.
7.4 Effect of ground movement
Vasilikis and Karamanos (2012) employed an advanced non-linear finite element method to
investigate the mechanical behaviours of a thin-walled lined pipe subjected to bending and to
determine the deformation of the lined pipe at wrinkles in the liner. 20-node brick elements
and shell elements were adopted for the host pipe and the liner respectively. From the numerical
results, it was found that the host pipe and the liner separated at the compression zone (liner
detaches due to host pipe bending) and consequently the liner buckled in the form of wrinkling.
It was also found that the wrinkling behaviour consisted of two bifurcations, with the first one
in a uniform wrinkling pattern and of small values of detachment and the second one of larger
values of wrinkling amplitude.
Bouziou (2015) numerically assessed the effect of transient ground deformation on a ductile
iron pipe with a circumferential crack or a weak joint. Field measured ground motion time
records were used as an input. Six deformation modes were investigated, including axial offset,
vertical offset, transverse horizontal offset, lateral rotation, vertical rotation and torsion.
Numerical results indicated that axial deformation is the most dominant type of deformation.
Therefore, to study the pull-out behaviour and investigate the failure mechanisms of CIPP lined
pipelines with circumferential cracks or weakened joints, Argyrou et al (2017) developed a
one-dimensional finite element model to simulate axial tension tests. Beam elements were
employed to model the liner and the pipe while non-linear springs were used to represent the
interface between the pipe and liner. For boundary conditions, one end of the numerical model
63
was fixed while tensile displacements were applied at the other end. It was found that numerical
results were in good agreement with the full-scale test results. Zhong (2015) developed a
number of simplified numerical models of buried ductile iron pipes with joints and lined with
InsituMain CIPP liners. The effect of properties of pipe-soil interaction elements, length and
mechanical properties of the joint, number of joints and pipe inertia were investigated. It was
found that pipe inertia and number of joints in the model have significant impact on the
response of the lined pipe with joints. It was also found that after rehabilitation with CIPP
liners, the probability of failure of the pipe under the maximum considered earthquake, is less
than 7%. This demonstrated the capability of the CIPP liner to significantly improve the pipe
performance.
7.5 Effect of creep
To study the long-term performance of a CIPP liner installed in deteriorated sewer pipes, Zhao
(1999) and Zhao et al. (2001) numerically investigated the creep-induced buckling of
constrained CIPP liners subjected to external pressure and the influence of geometric
parameters on liner buckling. Both the one-lobe (lobe is defined as a fold in the liner caused
by buckling) model and the two-lobe model were considered. In the author’s numerical models,
the compressive creep properties determined by Lin (1995) were employed. Based on the
numerical results, a model relating the critical liner buckling time to the applied external
pressure was developed.
Zhu and Hall (2001) investigated the change of contact conditions and stresses in a lined pipe
due to the liner deformation, for different levels of external pressure, material properties and
geometric parameters (e.g. ovalities, gaps, etc.). Flexural properties and compressive properties
determined by creep testing up to 3000 hours (Guice et al. 1994) were used for short-term and
long-term buckling, respectively. It was found that for thinner CIPP liners, the larger the
contact forces and areas are, the higher the enhancement factors are. Larger contact area
between the host pipe and the liner also leads to a short span of the lobe while the contact force
induced a reverse moment in the middle of the lobe. It was also found that the compressive
material properties seemed reasonable for liners with low ovalities and gaps with no axial
imperfections.
Guan et al. (2007) employed the results from the tensile creep testing up to 5000 hours, and
investigated the liner creep behaviour for a structural liner installed in a cast iron host pipe with
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circular defects. The diameter of the defects ranged from 102 to 203 mm. It was found that
creep resulted in an increase in the predicted displacement by 70% to 105% and the smaller the
defect size is, the higher the relative contribution of the creep effect to the displacement of the
liner is.
7.6 Summary
From literature, it was found that numerical modelling was employed to investigate the effect
of size and geometry of a defect on host pipe, material properties of the host pipe, liner
imperfections, ground movement and creep on the performance of the liner.
For the size and geometry of a defect on host pipe, different geometries of holes including
square, rectangle and ellipse (Guan et al. 2007), a missing segment of a host pipe (Brown et al.
2014) and circular surface patches (Shou and Chen 2018) were considered. Results showed
that rectangle and ellipse defect shapes with the long axes in the pipe longitudinal direction
show a significantly lower burst pressure than a circular defect. For the missing segment of a
host pipe, due to the greater strength and stiffness in the hoop direction, the maximum axial
stresses in the liner were reduced by approximately 27% and the tensile stress in the liner was
governed by the unconfined hoop stress at the maximum operational pressure. For the circular
surface patches, the internal pressure mainly controlled the stresses near the host pipe defect
area, while the displacements were mainly governed by surface loading. In addition, the stress
and displacement were significantly reduced at the damaged area by the installation of the CIPP
liner.
For material properties of the host pipes, two materials, namely, cast iron and PVC were studied
for the host pipe with defects of different sizes (Guan et al. 2007). Results showed that the
predicted liner burst pressures for cast iron and PVC host pipes are quite different when the
defect size is relatively small (e.g., the diameter of the defect was smaller than 100 mm), but
the burst pressure becomes similar as the defect size grows larger.
For the liner imperfections, the effect of liner longitudinal folds (Allouche et al. 2005;
Jaganathan et al. 2007), variation of liner thickness (Zhao 2003), localised wavy imperfections
(El-Sawy 2013) and a very small out of roundness (Rueda et al. 2012) were investigated on the
performance of the CIPP liners. For liner folds, it was found that when a longitudinal fold is
transverse to a hole in the host pipe wall, the burst pressure is the lowest. For variation of liner
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thickness, it was found that when the thickness only varies in the longitudinal direction, it was
found that the buckling pressure increases with the increase of the magnitude of the thickness
variation. When the thickness only varies in the circumferential direction, it was found that the
buckling pressure decreases with the increase of the magnitude of the thickness variation.
When the thickness varies in both directions, it was found that the buckling pressure decreases
with the increase of the frequency of thickness variation. For localised wavy imperfections,
results showed that an increase in the yield stress results in a decrease in the imperfection size
and liner thickness and convergence of the liner buckling to the elastic case with normalised
pressure close to unity. In addition, it was found that the liner’s normalised critical pressure
generally decreases with the increase of the imperfection angle. For a very small out of
roundness, three types of analyses, namely, static with hydrostatic elements, Riks general static
and general static, were considered and results showed that finite element models with
hydrostatic elements reasonably simulated the collapse of polymeric liners under external
pressure.
For ground movement, pipes subjected to ground motion time records (Bouziou 2015), axial
tension tests (Argyrou et al. 2017) and bending (Vasilikis and Karamanos 2012) were
numerically simulated. Numerical results from Argyrou et al. (2017) showed that the host pipe
and the liner separated at the compression zone (liner detaches due to host pipe bending) and
consequently the liner buckled in the form of wrinkling. Numerical results from Vasilikis and
Karamanos (2012) showed that pipe inertia and number of joints in the model have significant
impact on the response of the lined pipe with joints and that after rehabilitation with CIPP
liners, the probability of failure of the pipe under the maximum considered earthquake, is less
than 7%.
The behaviours of the lined pipes were also investigated numerically taking into account the
effect of creep (Zhu and Hall 2001; Guan et al. 2007). It was found that for thinner CIPP liners,
the larger the contact forces and areas are, the higher the enhancement factors are. Larger
contact area between the host pipe and the liner also leads to a short span of the lobe while the
contact force induced a reverse moment in the middle of the lobe. It was also found that creep
resulted in an increase in the predicted displacement by 70% to 105% and the smaller the defect
size is, the higher the relative contribution of the creep effect to the displacement of the liner
is.
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7.7 Gap identification
For the defects in the host pipe, only circular, elliptical and rectangular holes along the
longitudinal direction, a missing segment of the host pipe and circular surface patches were
studied on numerical modelling in literature. No inclined rectangular or elliptical holes were
considered. In addition, no numerical study has considered cracks in the host pipe.
For the study of liner imperfections, the effect of the variation of liner thickness and local wavy
imperfection were investigated for gravity pipes only rather than pressurised water pipes. In
addition, the effect of wrinkles in the circumferential direction, bulges, dimples and pinholes
etc. on the liner performance have not been investigated numerically. Apart from the above, no
studies have been carried out to investigate the behaviour of corroded host pipes with imperfect
liners subjected to ground movement and creep etc. Imperfections are deemed a major issue
and must be taken into consideration in a future design standard.
Further numerical studies for liner interaction would also be of benefit to a future standard.
8. Conclusions, research gaps, and future research 8.1 Conclusions
From literature the following conclusions were found:
• Limitations/Assumptions in current methodology for designing CIPP liners
o Defect geometry in host pipes (limited to circular or rectangular defects)
o Mainly designed for internal pressure
o Hydraulic loads for partially deteriorated host pipes and hydraulic, soil and live
loads for fully deteriorated host pipes.
o CIPP liner in the hole area considered as a uniformly pressurised plate with
fixed edges (not examining composite structure of lined pipe)
• Various types of defects in host pipes and imperfections in liners identified
• Different installation issues identified
• Short-term experimental studies on CIPP liners used for pressure pipes
o Most previous research on material properties of CIPP liners installed in sewer
pipes
67
o Limited number of studies on tensile and flexural properties (large variations
observed) of liners and bond strength between the host pipe and liner
o Most large-scale tests for Aqua-Pipe liners only
• Long-term pressure testing difficult to conduct due to end sealing problems
• Previous creep tests mostly on flexural creep and limited tests on tensile creep
• Previous research on standard creep testing focused on creep modulus reduction and
creep strain determination rather than strength reduction
• Accelerated creep testing (e.g., the SIM and SSM method) mainly applied to yarns,
HDPE and other thermoplastic materials but not to CIPP liners
• Numerical studies conducted to investigate the effects of defects in host pipe on the
behaviour of lined pipes
o Defects including circular, elliptical and rectangular holes along the
longitudinal direction, a missing segment of the host pipe and circular surface
patches
• Numerical studies conducted to investigate the effects of imperfections in CIPP liners
on the behaviour of lined pipes
o Effect of the variation of liner thickness and local wavy imperfection
investigated for liners in sewer pipes
o Effect of the longitudinal folds investigated for liners in pressure pipes
8.2 Research gaps
The following gaps were identified from literature:
• Aspects not considered in the current methodology for designing CIPP liners
o Inclined rectangular/elliptical defect, crack-like defect etc.
o Pressure transients, pressure induced thrust forces, Poisson effect, thermal
expansion effects and differential ground movement
o Effect of liner imperfections
• Limited short-term experimental studies on CIPP liners used for pressure pipes
o More research on tensile and flexural properties of liners and bond strength
between the host pipe and liner required for better statistics
o No negative pressure tests, shear and tensile tests found to be performed on
Starline, InsituMain and Saniline liners
o No pressure tests found on Starline and InsituMain liners
68
o No break tests were found on Saniline and InsituMain liners
• No standards available for the following tests
o Hole/gap spanning test, bending/shear tests, vacuum tests, pressure transient
tests
• Tensile creep tests only for up to 5,000 hours, and longer creep testing time is required
for better prediction of long-term behaviour of liners
• Not enough attention paid to the effect of creep on the strength reduction (e.g. tensile
and flexural) of the liners
• No studies on applying accelerated creep testing (e.g., the SIM and SSM method) for
CIPP liners
• Effect of inclined rectangular or elliptical holes and cracks in the host pipe on the
behaviours of lined pipes not investigated
• Effect of variation of liner thickness, local wavy imperfection, wrinkles in the
circumferential direction, bulges, dimples and pinholes etc. in the liner on the
behaviours of lined pipes not considered
• No studies carried out to investigate the behaviour of corroded host pipes with imperfect
liners subjected to ground movement and creep etc.
8.3 Future research
The following future research will be conducted at Monash University for the purpose of the
CRC-P project to address the gaps found in literature.
• Defects in host pipe
o Rank defects or installation issues
o Determine value of factor of safety for ranked defects
• Liner imperfections
o Quality control of liner variability
o Measurements of liner imperfections
• Properties of liners (short-term tensile and flexural)
o Cross-check data with testing on liners installed in Australia
o Liner suitability for different classes of liner
o Determine minimum strength requirement for Australian standards
• Adhesive strength needed for standard (Previously not included)
o Adhesion values for different surface types
69
o Adhesion properties are needed for initial liner installation (and tapping connections) and de-bonding, therefore a trade-off value of adhesion should be found for standards
o Check suitability for certain classes of liner (e.g. Class IV liner de-bonds from host pipe)
• Properties of liners (long-term tensile and flexural)
o Understand how the mechanical properties of liners change over time to quantify and predict strength reduction
o Extrapolate strength reduction factors for 20 and 50 years of testing o A thorough analysis of creep testing on different liners to determine whether a
50% stress reduction factors is generally applicable
• Pipe ring tests
o Determine crushing strength and flexibility of liner due to external loading o Examine de-bonding of the liner if host pipe is intact
• Vacuum test
o Examine whether de-bonding is an issue o Check whether inherent ring stiffness is present
• Hole or gap spanning tests
o Check the ability of liner to contain pressure under a partially deteriorated host
pipe
• Pipe bend/shear tests
o Determine bending/shear strength of pipe liner bonded to host pipe under
internal pressure
o Verification of liner de-bonding over a larger pipe length under internal pressure
• Laboratory short-term hydrostatic pipe burst test
o Determine hydraulic capacity of the liner with different defect sizes
o Test if liner can sustain pressure over a certain period
• Laboratory pressure transient tests
o Determine hydraulic capacity of the liner with different defects o Test if liner is able to sustain cyclic transient pressures o Conduct 10,000 cycles with defect to determine if fracture propagation of the
liner is an issue
• Accelerated creep/long-term strength tests
• Numerical modelling of lined deteriorated pipes
o Compare numerical modelling results with laboratory testing results for
validation of numerical models
70
o Predict liner performance under various external and internal loading conditions
o Determine how the defects/imperfections in both the host pipe and liner affect
the strength of the liner
o Regression models will be derived from numerical models
o Results will be implemented in the decision tool for appropriate liner selection
9. Acknowledgements
The smart linings for pipe infrastructure project (CRC-P) is a collaborative project funded by
the Australian Government Business Cooperative Research Centres Program, Water Services
Association of Australia, Coliban Region Water Corporation, Hunter Water Corporation, Icon
Water Limited, Melbourne Water Corporation, South Australian Water Corporation, South
East Water Corporation, Sydney Water Corporation, Northern SEQ Distributor – Retailer
Authority, Water Corporation, Bisley & Co PTY LTD, Insituform Pacific Pty. Limited,
Kerneos Australia Pty Ltd, Parchem Construction Supplies Pty Ltd, Abergeldie Watertech Pty
Ltd, Interflow Pty Ltd, Ventia Pty Ltd, Metropolitan Restorations Pty Ltd, ITS/Downer
Pipetech Pty Ltd, Monadelphous Group Limited, UKWIR, Water Research Foundation, Water
Environment and Research Foundation, Calucem, Milliken, and Sanexen Environmental
Services.
Research Partners are Monash University, University of Sydney and University of Technology
Sydney (UTS).
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