Materials and Structures (2021)54:51
https://doi.org/10.1617/s11527-021-01634-9(0123456789().,-volV)( 01234567
89().,-volV)

ORIGINAL ARTICLE

Prestressing low clinker structural concrete elements
by ultra-high modulus carbon fibre reinforced polymer
tendons
Tobias Dominik Lämmlein . Janis Justs . Giovanni Pietro Terrasi .
Pietro Lura

Received: 15 July 2020 / Accepted: 17 January 2021
 The Author(s) 2021

Abstract The combination of low clinker highperformance concrete (LCHPC) and ultra-high modulus (UHM) carbon fibre reinforced polymer (CFRP)
tendons was recently proposed for prestressed structural elements. The 70% reduction in cement content
resulting in limited creep and shrinkage of the LCHPC
in comparison to a conventional high-performance
concrete (HPC) and the very high UHM-CFRP tendon
stiffness ([ 509 GPa) were expected to impact themechanical behaviour of such structures. This study
focuses on the behaviour of 3 m-long beam specimens
during prestressing, concrete hardening and in 4 pointbending experiments. Fibre optic sensors were implemented inside the CFRP tendons to measure strain
during those stages and a digital image correlation
system was employed to monitor the 4-point-bending
T. D. Lämmlein (&)  J. Justs  G. P. Terrasi  P. LuraEmpa, Swiss Federal Laboratories for Materials Scienceand Technology, Überlandstrasse 129, 8600 Dübendorf,
Switzerlande-mail: Tobias.Laemmlein@empa.ch
G. P. Terrasi
e-mail: Giovanni.Terrasi@empa.ch
T. D. Lämmlein  P. Lura
Institute for Building Materials (IfB), ETH Zurich,
8092 Zurich, Switzerland
G. P. Terrasi
Institute for Infrastructure and Environment, School ofEngineering, University of Edinburgh,
Edinburgh EH9 3FB, UK

tests. After 28 days, the LCHPC recipe, despite a 70%
cement reduction and much smaller environmental
footprint, did not show measurable differences in theprestress loss behaviour in comparison to a conventional HPC. The UHM-CFRP prestressing tendons,because of their stiffness, showed both higher prestress losses of around 40% and on average a nearly
doubled prestress transfer length. However, they
increased the beam‘s maximum load-bearing capacity
by 21% and showed 47% less deflection at failure incomparison to beams prestressed with the standard
modulus (UTS)-CFRP tendons.
Keywords Prestressed concrete  Low-clinker-highperformance-concrete (LCHPC)  Carbon fiber 
Optical fiber bragg grating  Prestress-losses  Prestresstransfer  Bending behavior

1 Introduction
Sustainable and innovative designs in construction canmake a significant contribution for improving thecarbon footprint of structures in the future. In terms ofstructural design, there are two main approaches totackle this challenge. The first approach consists in
reducing the environmental impact of the employedmaterials. The second approach is based on optimizing

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the mechanical performance at the structural level andhence the effectiveness in the use of materials.
Already in the 1990s, one first step for improving
the structural design of building facades was thedevelopment of slender carbon fibre reinforced polymer (CFRP) prestressed elements [1, 2]. However, fornearly two decades, the space for further developments in this field was constricted by both the highprice of the materials and by existing standards.
Currently, standard modulus CFRP costs about 20 €/kg. Ultra-high modulus (UHM) carbon fibre reinforced polymers have become recently more affordable with a cost of about 50 €/kg [3]. In addition, newguidelines [4, 5] are allowing the development ofnovel low clinker high performance concretes(LCHPC) with clinker replacement levels of up to
70% [6]. The combination of both materials results ina very promising composite system that satisfies boththe sustainability aspects and the technical improvement objectives.
Prestressing of concrete became technically feasible with the invention of high strength steels and was
first realized by Freysssinet in the 1920s [7]. By
counteracting tensile stresses in concrete, the prestressing technique allowed the manufacturing oflighter structural elements at much lower cost. Laterin the 1970–1980 s, fibre reinforced polymers werefirst brought to use in concrete structures [8]. Due totheir excellent strength, their low weight and theircorrosion resistance, these materials gained international attention and were widely investigated forstrengthening, retrofitting and prestressing of structures. In particular CFRP, having the highest strengthand stiffness among all fibre reinforced polymers, canbe seen as the technically ideal material for prestressing applications [9]. In contrast to glass-fibre andaramid-fibre reinforced polymers, CFRP is not subjected to stress corrosion [10] or creep [9, 11],
respectively. Since the invention of the prestressing
technique, the load deflection behaviour of a prestressed concrete structure was linked with the level ofprestress and hence dependent thereon. In consequence, it was crucial for the designers and engineers
to assess the amount of prestress loss over time, which
needed to be considered in their designs. First
investigated by Magnel [7, 12], this finally resulted
in numerous codes and recommendations, e.g.
[13–16]. In the case of internal prestressing by CFRP
tendons1 using the pre-tensioning technique, the main

Materials and Structures (2021)54:51

contributors to prestress losses can be seen as elastic
shortening, shrinkage and creep of concrete. The
measurement and monitoring of prestress losses was
performed by different approaches. In [17], Maaskant
et al. implemented fibre-optic Bragg grating (FBG)
sensors on steel and CFRP prestressing tendons tomonitor the prestress losses. The losses were found to
be 25% lower for CFRP tendons compared to steel
tendons, mostly due to the lower stiffness of the CFRP.
In a later study [18], detachable mechanical strain
gauges were placed at a 50 mm distance on the
sidewall of prismatic concrete specimens, to evaluateprestress losses without disturbing the bond between
tendon and concrete. In combination with prestressforce measurements and additionally performed
shrinkage tests, they were able to show that shrinkage,
creep and elastic shortening contribute each withsignificant amounts to prestress losses. Total losses
ranged from 25% up to 60% over one year, dependent
on specimen size and concrete recipe. Direct application of strain gauges on the surface of the tendons was
also employed [19], however neglecting bond disturbance. With this technique, they were able to record aprestress development up to 600 MPa in a selfprestressed configuration which was driven by an
expanding concrete recipe. Recently, a prestressed
concrete strand with an integrated FBG sensor to allow
prestress measurements in concrete structures wasdeveloped [20]. The center wire of the strand was
exchanged with a CFRP wire and a FBG sensor wasinserted into the CFRP wire during production. They
were able to measure strains up to 14,000 le, which islarger than the typically design prestress of a conventional steel strand.
In pre-tensioned elements the prestress needs acertain distance starting from a structure’s boundary to
be transferred by bond to the concrete. The distanceneeded to reach the fully prestressed section in aprestressed concrete element is commonly referred toas the prestress transfer length. Knowing precisely
where a structural component is fully prestressed andhence external loads could safely be introduced is
indispensable. In [21], a pioneering work, Bruggeling
outlined all major steps to describe transmission ofprestress by bond. Up to that paper, no realistic model1

Note: The prestressing elements called ‘‘tendons’’ in thispaper, see [3, 28 37], are also commonly called ‘‘strands’’ in theliterature‘‘.

Materials and Structures (2021)54:51

had been developed that was able to describe thecontext of prestress transfer as a whole, neither has thistheoretical feat been accomplished since. On the other
hand, numerous experimental studies investigating theprestress transfer length of FRP tendons were conducted in the last 20 years. The main outcome in [22],a study of the transmission mechanism of prestress toconcrete by FRP, was that the transfer length is mainly
affected by the surface characteristics and the size ofthe tendons. Furthermore, in [22], it was shown thatthe mechanism of load transfer is different in FRP
compared to classic steel tendons. In this study, FRP‘stransfer was seen related to adhesion and mechanicalinterlock, while steel tendons were subjected totransfer by mostly friction. A study about the force
transfer of FRP bars in concrete [23] showed that asand-coated FRP bar had a stress transmission length
limited to only 16Ø and proved that this length was
smaller in comparison to steel tendons.
The tensile behaviour of CFRP is linearly elastic
until failure, in contrast to the elastic–plastic responseof steel. It is therefore to be expected that thebrittleness of the CFRP will affect the structural
behaviour of a prestressed concrete structure. The
structural behaviour of precast prestressed beams,
façade elements and slabs requires further investigation and the development of new guidelines. Several
institutions around the world were involved in thisprocess for FRP materials since the 1980s. Theirfindings and approaches have first been reviewed in[24] and more comprehensively in [25]. Both reviews
commented that the used FRPs had a lower elastic
modulus in comparison to steel and hence they would
increase the deflection and the crack width in loaded
concrete structures. Additionally, in a study about therational use of advanced composites in concrete, it wasstated that the structural use of FRPs is generally
feasible [26]. Later, Bakis concluded that FRP applications would probably be restricted to cases where
their properties of low weight, high strength or
corrosion resistance are crucial [25].
The short-term and long-term flexural behaviour of
FRP prestressed concrete elements were studied
experimentally and modelled analytically [1, 27–32].
Recently, Rezazadeh et al. proposed a new simplified
approach to predict the flexural behaviour of CFRPprestressed, simply-supported beams [33]. This model
was developed in the scope of flexural strengthening
with externally bonded and prestressed strips. The newPage 3 of 18

51

method, based on strain compatibility and on theprinciple of static equilibrium, showed a good predictive performance for the load deflection behaviour. Allof the discussed studies used CFRP tendons withlongitudinal stiffness’s lower than 200 GPa[27–30, 33, 34]. The new UHM-CFRP tendons
considered in this study are substantially stiffer thanclassic CFRP materials and even stiffer than steel. In
this context standard modulus carbon fibres are
typically referred to have a stiffness in the range of
220–240 GPa and carbon fibres with UHM properties
a stiffness of more than 520 GPa [35]. Their corresponding ultimate tensile strength lies in the range
from 2050 to 3790 MPa for standard modulus and
1380–2400 MPa for UHM-CFRP fibres. However, thelatest developments of UHM-fibres are even exceeding this strength. It could be anticipated that thisdifference in stiffness will have significant influenceson the structural behaviour. It must also be remarked
that the reduced clinker content in the novel LCHPCs
would cause faster carbonation and hence would be alimitation for the use of classic prestressing steel. The
proposed UHM-CFRP tendons, not unlike other CFRP
materials, are practically immune to corrosion andhence ideal for the proposed implementation in
LCHPCs.
The goal of this work was to provide a firststructural characterization of LCHPC beams prestressed with UHM-CFRP prestressing tendons. To
support the design of efficient and sustainable structures, the effects of the new LCHPCs and of the UHM
CFRP tendons on the prestress behaviour were ofparticular interest. This study identifies the amounts ofeach individual contributor to the total prestress loss,
precisely measures the prestress transfer length of the
CFRP tendons and discusses approaches for improving the ultimate load bearing capacity of CFRPprestressed beam elements.

2 Materials and methods
New materials for prestressing of structural concreteelements, UHM-CFRP and LCHPCs introduced in2016 [3] and 2019 [6] respectively, were tested toobserve their behaviour during prestressing and during
4-point bending tests of such elements. In addition, thestructural behaviour of these beams was further
characterized. The related materials, their

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characterization methods and the experimental program are described in this section.
2.1 Raw materials and composites
A high performance self-compacting concrete(HPSCC) was selected to investigate the influence of
the CFRP prestressing tendon‘s stiffness on the
structural performance of prestressed concrete beams.
This concrete recipe with a water to binder ratio (w/b)
of 0.35, named C1, contained 443.0 kg/m3 of ordinary
Portland cement CEM I 52.5 R, 120.0 kg/m3 fly ash,20.0 kg/m3 silica fume, 1107 kg/m3 sand 0/4, 487 kg/
m3 gravel 4/8 and 5.2 kg/m3 superplasticizer (SP).
This is the same recipe that was recently employed as areference mix during the development of LCHPCs [6],based on a HPC used by the Swiss precast industry in
CFRP prestressed structural elements [36]. In addition, to evaluate the potential of LCHPCs for prestressing applications, the recipe C4 (w/b = 0.17) waschosen from three recently developed LCHPCs [6]. In
this recipe (now containing only 134.3 kg/m3 of
cement), large amounts of cement were substituted
with 511.7 kg/m3 limestone filler. Further, the amountof silica fume was increased to 67.2 kg/m3. The
amount of sand and gravel was the same as in C1. The
clinker reduction between C4 and C1 was 70%; both
mixtures were presented in detail recently [6].
All concrete compositions were mixed in a 250-l
Eirich RV11 mixer. For the recipe C1, all drycomponents were tumbled for 60 s, afterwards water
was added and finally the concrete was mixed for
120 s until a uniform mixture with good flow properties was achieved. The LCHPC mixture C4 wasmixed in a slightly adapted manner. Tumbling of thedry components was performed but with a reduced
amount of only 100 kg sand 0/4. Then water wasadded and the recipe was mixed until the superplasticizer started functioning. Finally the remaining
aggregates of sand 0/4 and gravel 4/8 were added
and the recipe was mixed for another 120 s until auniform mass with good flow properties was achieved.
Air void content and density, were measured
according to EN 12350 (2009) directly after mixing,
see Table 1. Concrete flow was measured in [6] with
72.5 cm for the C1 recipe and 57.5 cm for the C4
recipe. Subsequently after mixing a total of nine
150 9 150 9 150 mm3 cube specimens and five

Materials and Structures (2021)54:51

beam specimens were cast from one mix for eachconcrete/CFRP combination.
The compressive strength of the concretes wastested on three 150 9 150 9 150 mm3 cubes after 3
and 28 days, following standard EN 12390 (2009). In
addition, the static E-modulus of the concretes wasmeasured after 28 days on cylinder specimens(d = 50 mm, h = 150 mm), following the standard
EN 12390 (2013) Method B. Until testing, the samples were stored in a climatic room at 20 C ± 0.3 C
and RH [ 95%. The storage conditions of the beamspecimens are specified in the general part of Sect. 2.2.
The static E-modulus of LCHPC C4 was
41.7 ± 0.2 GPa and the one of the HPC C1 was
35.1 ± 0.8 GPa. Further test results of all investigatedconcretes can be found in Table 1. Shrinkage andcreep tests on the respective concretes were not
performed separately during this study. Those results
were gained previously in [6] generally following
SIA 262 at 20 C ± 0.3 C and at 70% RH ± 3%
RH. In [6], after 28 days, shrinkage was measured at- 0.38% for the HPC C1 and – 0.17% for the
LCHPC C4; the 28 day creep measurements showed- 0.61% and - 0.41% respectively. In contrast to
SIA 262, the creep samples were loaded in a sequenceof increasing compressive stress applied at 3, 7 and
28 days after casting; at each loading step, the total
stress was changed and corresponded to 20% of therespective compressive strength at this time point.
Details can be found in [6].
In this study, two significantly different CFRPtendon materials were investigated both with a fiber
volume content of 60%. The first tendon type was astandard-modulus CFRP material (UTS-CFRP) andconsisted of Tenax-J/E-UTS50 carbon fibres (Teijin
Carbon Europe GmbH, Germany) in combination witha Huntsman XB3515/AD5021 hot-melt epoxy system.
The second tendon type was made out of Mitsubishi
DIALEDTM K13916 ultra high modulus (UHM)
carbon fibres (Mitsubishi Rayon Co. Ltd., Japan) also
in combination with the Huntsman XB3515/AD5021
hot-melt epoxy system; hereafter named UHM-CFRP.
The tendon materials were characterized in tensile
tests following the standard EN 2561. The longitudinal elastic modulus of the UTS-CFRP tendon was E11
of 145 ± 3 GPa and the one of the UHM-CFRP
tendon was E11 of 509 ± 13 GPa [3]. The corresponding tensile strength values can be found in
Table 4. Both tendon types were produced by

Materials and Structures (2021)54:51

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51

Table 1 Properties of Concretes
Concrete recipe

Compressive
Strength
3 daysCompressive
Strength
28 daysCompressive
E-Modulus
3 daysCompressive
E-Modulus
28 daysTensileStrength
28 days

(MPa)

(GPa)

(GPa)

(GPa)

(MPa)

SD

SD

SD

SD

SD

Density

Air void content(kg/m3)(%)

C1(HPC)-reference*

56.9

0.5

85.7

1.1

29.0

0.9

35.4

0.4

6.6

0.8

2307

3.6

C1 (HPC)-UTS-Mix

61.8

0.3

87.5

1.6

n.a

n.a

35.2

0.6

n.a

n.a

2272

3.5

C1 (HPC)-UHM-Mix

62.8

0.3

87.9

1.8

n.a

n.a

35.1

0.8

n.a

n.a

2274

3.9

C4 (LCHPC)reference*

35.1

0.1

83.3

1.0

26.8

0.2

38.8

0.6

4.7

0.6

n.a

n.a

C4 (LCHPC)-UHMMix

32.9

0.5

88.3

1.2

n.a

n.a

41.7

0.2

n.a

n.a

2368*

3.2*

*Data equivalent to[6]n.a. means not available
SD means standard deviation
Air void content and density following EN 12350

Carbolink AG (Fehraltdorf, Switzerland) in a tapelaying method with a cylindrical surface and adiameter of d = 5.3 mm. After curing, the tendons
were coated with quartz sand particles (63–500 lm) toenhance the bond between the tendons and theconcrete. The chosen sand coating showed its effectiveness recently not only due to its strong bond to
concrete but also due to its bond durability [28, 37].
The sand-coated UHM-CFRP tendon was the same
tendon type as previously used in [3, 36] but produced
in another batch.
2.2 Prestressed concrete beam elements
To evaluate the structural behaviour of UHM-CFRP
prestressed LCHPC elements, a slender beam geometry was chosen. The beam length was set to3000 mm, with a height of h = 157.5 mm and a widthof w = 90 mm, see Fig. 1. This geometry followed,
albeit scaled down to , an earlier work in the field of‘‘Pretensioned prestressed concrete members’’ by
Nanni and Tanigaki [38]. The prestressing tendons
were positioned at a distance of 22.5 mm to the bottom
and the sidewalls of the beam respectively. The
resulting low concrete cover and the chosen prestresslevel of 800 MPa inside the CFRP prestressing
tendons were based on recent literature [39]. In
contrast to practical prestressed concrete design, in
this study, the prestress level of 800 MPa was not

Fig. 1 Cross section of CFRP-prestressed concrete beamadapted according to the different CFRP or concretematerial properties. Having the same initial prestressenabled a clear and easy comparison between allinvestigated material configurations.
The beam elements were produced at Empa by theaid of a specially-designed prestressing rig. This rig
allowed the production of up to five beam specimens
and the application of the prestressing force to all
employed prestressing tendons at once by using two

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hydraulic cylinders. In addition, to reach a consistentprestress level of 800 MPa ± 5%, each tendon could
be adjusted individually by a thread mechanism. The
initial prestress level was monitored for each tendon
by the aid of a strain gauge which was applied in thefree length of the tendon between the formwork andthe tendon clamping mechanism. The prestress was
released three days after casting. Afterwards the
samples were stored in a climate chamber at
20 C ± 0.3 C and 90% ± 3% RH until 21 days
after production. Then the beams were transported into
the mechanical testing laboratory and stored until
28 days after casting in controlled climate conditionsof 23 C ± 2 C and 50% ± 5% RH.
In total, 15 concrete beam elements were cast, 5
beams each for the following combinations; UHMCFRP and C1, UTS-CFRP and C1, UHM-CFRP and
C4. All three sets consisting of 5 beams were produced
from one single batch of concrete.
2.3 Optical fibre strain measurements
The tape-laying production method of the CFRP
tendons allowed for the implementation of fibre optic
sensors in the centre of the CFRP prestressing tendon
along the tendon‘s longitudinal axis. The very thinfibre optic sensor (d = 125 lm) had no significant
influence on the diameter of the CFRP prestressing
tendon but allowed strain measurements inside thetendon without affecting the tendon‘s sand-coatedsurface. Each sensing point inside the fibre optic
sensor was represented by a 5 mm-long fibre Bragg
grating (FBG). A FBG is in principle a periodic
microstructure with a varying refractive index which
can be written into the core of the optical fibre byintense ultraviolet light. When strained, the FBG
elongates, the fibre index changes and hence thereflected Bragg wavelength of the FBG shifts. The
measurement of this shift can be analysed andcorrespondingly the strain at the position of the FBG
can be calculated [40]. A detailed summary of the
FBG measurement principle, their possible applications and variations can be found in [41].
The fibre optic sensors used in this work were DTGA3A4-A01 fibres produced by FBGS International NV(Geel, Belgium) and supplied by Com&Sens bvba
(Eke, Belgium). To preserve the sensitive optical fibre
from damage, a thin Teflon tube of d = 0.9 mm
covered the optical fibre as far as 0.32 m ahead of theMaterials and Structures (2021)54:51

first sensing point. In addition, to safely protect the
optical fibre when outside the CFRP tendon and up toreaching the connector to the measurement device, arobust Teflon tube of d = 3 mm was additionallyemployed.
The available 15 FBG sensing points in the opticalfibre were positioned to be able to best record theprestress transfer length. Their detailed positions
along the 3 m-long concrete beam specimen can befound in Table 2. The final positioning of the sensing
points in the respective prestressed concrete beam
element were determined by the aid of a reference
marking, which was placed on the 3-mm protective
Teflon between the end of the concrete beam and theoptical fibre connector. Based on the distance measured between the reference point and the concrete
beam and the condition of zero stress in the tendon
outside the beam after prestress release, the exact
position of the first FBG was evaluated and used as
reference during the analysis. Before casting, theoptical fibres were connected to an FBG SCAN 804D
high precision measurement device and the reference
zero strain level was set in the analysis software
Illumisense V2.2.
The prestress transfer length, see Table 2, wasmeasured based on the sensing points of Group A andthe measurement of the prestress level was based on
the sensing points of Group B. As proposed inliterature [42], the prestress transfer length calculation
was based on reaching 95% of the full prestressmeasured in Group B. The final prestress transfer
length was calculated by linear interpolation between
two sensing points of Group A where the 95%
criterion was reached. The prestress level after casting
was calculated for all time points and beam configurations in relation to the first strain measurement taken
three days after casting right before the prestressrelease. This first measurement was considered tocorrespond to 100% prestress and the subsequent
prestress levels were calculated based on this point.
2.4 Digital image correlation
A LIMESS-Q400 two-camera digital image correlation system (LIMESS Messtechnik u. Software
GmbH, Germany) was set up for monitoring the beam
elements during loading. In combination with theanalysis software Istra4D (Dantec Dynamics A/S,
Denmark), this arrangement allowed measuring theMaterials and Structures (2021)54:51

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51

Table 2 Specified FBG sensing point position along the concrete beamSensing point #

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

Position along beam length (mm)

0

12

24

36

60

84

108

156

204

300

996

1248

1752

2004

2760

Group

A

A

A

A

A

A

A

A

A

A

B

B

B

B

A

concrete beam‘s deflection and the strain distribution
on the vertical face in the beam‘s middle third.
The two cameras, each with a 100 sensor and 6 MP
resolution, were mounted in a stereoscopic setup on
one tripod and placed approximately 1.8 m in front of
the experimental setup. The lenses of the cameras had
a focal length of 16 mm. The field of view (FOV) in
this arrangement was about 1.5 m wide. The camera
set up was calibrated with an AL_35_WMB_9 9 9calibration target using the implemented procedure in
the analysis software Istra 4D. On the vertical surfaceof the concrete beams, facing the cameras, a randomly-distributed contrast pattern (speckle pattern)
was applied. Pictures were taken during the 4 point
bending experiments every 2 s. The local surface
pattern was automatically assigned in the camera
images and the strain distribution as well as the beam‘s
deflection could be evaluated.
2.5 Four point bending experimental setup
To investigate the influence of a LCHPC and a UHMCFRP prestressing tendon on the load bearing capacityof prestressed concrete beams, a simple 4-point
bending test configuration was chosen. The distance
between the beam support points was set to 2500 mm.
This allowed an overhang of 250 mm on each side of
the beam and avoided interaction between the prestress transfer area and the loaded beam section. The
load was applied by a servo hydraulic cylinder andtransferred to the beam by means of a spreader beam.
The range of constant bending moment was set to
840 mm in this configuration. All contact points to theconcrete beam were steel cylinders with a diameter ofd = 50 mm. An overview of this 4-point bending set
up is shown in Fig. 2. Moreover, the mid beam
deflection was measured by a linear variable differential transformer (LVDT) of type W20TK HBM(Hottinger Baldewin GmbH (HBM), Germany) tovalidate the DIC measurement. Results shown by the

LVDT and the DIC showed negligibly small differences. Hence, possible deformations of the setup,concrete crushing or small geometrical imperfections
did not affect the values recorded. All beams were
loaded continuously at a servo hydraulic cylinder
velocity of 1 mm/min until failure.

3 Results
3.1 Prestress transfer length and loss
The first strain measurements inside the CFRP
prestressing tendons were taken three days after
casting, right before and after the prestress release.
The measurement was repeated for all beams at 7, 14,
21 and 28 days after casting. One additional measurement was performed for C4-UHM-CFRP beams
4 days after casting.
All five beams of type C1-UTS-CFRP were
equipped with one sensing and one non-sensing CFRP
tendon. One sensing CFRP tendon was damaged
during transport after the 21 day measurement. In the
beam configurations of type C1-UHM-CFRP and C4UHM-CFRP, four of the five beams were initiallyequipped with one sensing and one non-sensing CFRP
tendon. In these groups of beams, the remaining beam
was equipped with two non-sensing CFRP tendons.
However, for the C4-UHM-CFRP beams, the very
brittle material behaviour of the UHM-CFRP led to theloss of one sensor cable during demolding and the lossof another during transport after the 21 day
measurement.
After 28 days, the average prestress transfer length
in the beams C1-UHM-CFRP was found to be
181 mm with a standard error of the mean (SEM)
of ± 28 mm. The same behaviour was found for thebeams C4-UHM-CFRP with a prestress transfer lengthof 148 mm and a SEM of ± 18 mm. The beams using
a standard modulus CFRP tendon, C1-UTS-CFRP

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Materials and Structures (2021)54:51

Fig. 2 Four point bending test configuration: Overview of mounted sample in experimental set-up, span 2500 mm, load span 840 mm
and shear span 830 mm

showed a significantly shorter prestress transfer lengthof only 93 mm and a much smaller SEM of ± 4 mm.
The difference in the SEM between beams prestressedwith UHM-CFRP tendons and beams prestressed with
UTS-CFRP tendons can be explained by the huge
difference in the tendon‘s longitudinal stiffness. The
stiff tendons react much more sensitive to material
inhomogeneities or imperfections and hence show amuch larger scatter. At selected time points between
prestress release and until 28 days after casting, nosignificant change in the prestress transfer length
could be observed for any of the investigated beam
types. The detailed results of the prestress transfer
length measurement can be found in Fig. 3. In thisFigure, the obvious drop in the prestress transfer
length for C4-UHM beams after 28 days was causedby the loss of one sensing tendon. As a consequence,
the last calculation of the averaged prestress transfer
length for C4-UHM-CFRP beams, see Fig. 3, wasbased only on two beams.
The initial prestress level after release, see Fig. 4,
was on average 75% for C1-UHM-CFRP beams, 96%
for C1-UTS-CFRP beams and 75% for C4-UHMCFRP beams, respectively. The initial camber of thebeams after the prestress release was calculated basedon the remaining prestress force to be 0.75 mm for C1UHM CFRP prestressed beams, 0.7 mm for C4-UHMCFRP prestressed beams and 1.32 mm for C1-UTSCFRP prestressed beams. Afterwards, the prestresslevel of all investigated beams showed a clear decreasein the first 28 days. The total loss (42% of the initialprestress) was most pronounced in the beams C4UHM-CFRP. The C1-UHM-CFRP beams showedalso a high loss of 40%, while the C1-UTS-CFRP
beams showed significantly smaller losses of only
11% when measured 28 days after casting. A detailed
overview of the prestress level and its behavior in thefirst 28 days for each configuration can be found in
Fig. 4. The prestress development after 28 days andthe position of the prestress transfer length in the endzone of the beam is given for completeness in Fig. 5,
the shown prestress development is representative forall time points between 3 and 28 days. In addition, thestresses inside the concrete after 28 days related toprestressing were calculated based on the resultspresented above and by using the adapted model of[33]. In the compressive zone C1-UTS-CFRP beams
were compressed on average by 0.020%, C4-UHMCFRP beams by 0.014% and C1-UHM-CFRP beams
also by 0.014%, the stresses were 5.8 MPa, 3.75 MPa
and 4.06 MPa respectively.
3.2 Concrete beam deflection and maximum load
bearing capacity
The aim of this experimental program was to investigate the influence of UHM-CFRP prestressing tendons and LCHPC on the structural behaviour ofMaterials and Structures (2021)54:51

Fig. 3 Prestress transfer length development in the first 28 days
after casting. The bars show the average value of the beams
investigated and their standard deviations. The drop in prestress

Page 9 of 18

51

transfer length for the C4-UHM-CFRP beams after 28 days iscaused by the loss of one sensing tendon after the 21 day
measurement

Fig. 4 Prestress level development in the first 28 days after casting. The bars show the average value of the beams investigated andtheir standard deviations

prestressed concrete elements. For this purpose, three
beams of type C1-UHM-CFRP and C4-UHM-CFRP
as well as two beams of type C1-UTS-CFRP were
tested in 4-point bending tests. Two of the beams with
C1-UHM-CFRP tendons and one of the beams for thetwo other configurations were equipped with an FBGsensors. The remaining beams were stored for longterm outdoor tests which were not in the focus of thisstudy.
A similar deflection behaviour was observed in all
beam types in the first branch of loading, with nearly
constant slope of about 3.40 kN/mm for all beams. In
the second branch, beam types C1-UHM-CFRP and
C4-UHM-CFRP showed a load increase rate of

0.88 kN/mm and 0.97 kN/mm respectively. The beam
type C1-UTS-CFRP deflected in a softer manner andhad a load increase rate of 0.23 kN/mm, see Fig. 6 A.
In the transition zone between the first branch and thesecond branch, the rate of increase of the loaddecreased progressively. The tendon draw-in was
initially measured with the same setup as employed in[36] up to 70% of the ultimate load but no draw-in
could be recorded. The measurement system had to be
removed due to safety reasons thereafter. To monitor
transfer and bond along the tendon’s length the strain
measurement by FBG inside the tendons wasconsidered.

51

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Materials and Structures (2021)54:51

Fig. 5 Stress inside prestressing tendons as measured by the
FBG-sensors. Orange indicates the new C4-UHM-CFRP beam
type, dark grey the C1-UHM-CFRP beams and light grey thecurrent industry standard C1-UTS-CFRP type. a shows theresults 28 days after casting, and at ultimate load during 4-point

bending tests. b Shows the averaged prestress development in
the tendon of the respective beam-tendon combination 28 days
after casting, including their transfer length (PSL) as verticallines

Fig. 6 Results of 4p-bending tests performed, 28 days after
casting, on CFRP prestressed beam specimens. Orange indicatesthe new C4-UHM-CFRP beam type, dark grey the C1-UHMCFRP beams and light grey the current industry standard C1UTS-CFRP type. All beams were initially prestressed to thesame level of 800 MPa. The average bi-linear load deflection
behaviour for the different beams is shown in a, together with
their standard deviations. The corresponding maximum load
bearing capacity is visualized in b, also together with theirrespective standard deviations

All beams failed by a sudden and complete
breakdown due to CFRP tendon rupture; at the beams‘
end faces no tendon slippage could be observed. The
averaged maximum failure load is shown for all beam
types in Fig. 6 B. The C1-UTS-CFRP beams had
clearly the lowest load carrying capacity of 21.68 kN.
The beams C1-UHM-CFRP and C4-UHM-CFRP
reached on average maximum loads of 26.31 kN and26.37 kN respectively. The corresponding maximum
beam deflections at the middle of the span were
measured for C1-UTS-CFRP beams with an averageof 37.85 mm. The beams C1-UHM-CFRP and C4UHM-CFRP deflected less and reached an average of
20.76 mm and 19.18 mm respectively. At the point of
failure, the ultimate strain inside the CFRP prestressing tendons was 1.06% in the case of UTS-CFRP

Materials and Structures (2021)54:51

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51

tendons and on average 0.38% in the case of UHMCFRP tendons. The main results of the 4-point
bending tests for each particular beam tested aresummarized in Table 5.
Furthermore, the strain distribution on the verticalface of the concrete beams was measured with the DIC
system. Typical results for each type of beam areshown in Fig. 7. In this visualization, it is clearlyvisible that the cracks in the C1-UTS-CFRP beams
open more in comparison to both UHM-CFRP
prestressed beam types. The beams which are notshown in this graph had very similar strain patterns
within their respective group of beams.

Fig. 7 Visualization of the surface strain pattern along thehorizontal direction of the beam‘s vertical face at selected load
levels. Level A corresponds to a load of 5 kN, level B to thecracking load and level C to 20 kN. The left column represents abeam of type C1-UHM-CFRP, the middle column an industry
standard type C1-UTS-CFRP and the right column the new C4-

UHM-CFRP beam type. The distance between the cracks in thepattern of the different beam types was similar; all showed about
6 cracks in the same field of view. The crack width was visuallylarger for beams prestressed with C1-UTS CFRP tendons. A
detailed crack width analysis was not performed

4 Discussion
4.1 Effects of UHM-CFRP tendons in prestressed
concrete
The implemented UHM-CFRP tendons were 3.5 times
stiffer in comparison to their UTS CFRP counterparts.

51

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Not surprisingly, their response to the beam‘s strainchanges, due to shrinkage and creep after the prestressrelease, were found to be more significant and resulted
in much lower tendon stress levels. In addition, themeasurement of the prestress transfer length and theprestress level revealed a large scatter for beams
prestressed with UHM-CFRP tendons; see results in
Figs. 3 and 4. Both of these values are generally
dependent on the effective strain distributions inside
the beams that are reached after prestress release(when elastic shortening of the concrete happens) but
also over time (due to creep and shrinkage of theconcrete). In this context, local concrete material
variations, geometrical beam imperfections and prestress tendon positioning tolerances seem likely toaffect the prestress-related parameters in UHM-CFRP
prestressed beams more significantly than in UTSCFRP prestressed beams. Due to the higher longitudinal stiffness of UHM-CFRP, this effect is proposed
to be resulting in the larger scatter found in theexperiments. In other words, beams prestressed with
UHM-CFRP can be expected to be more sensitive intheir resulting prestress force and this uncertainty
should be considered when designing such beams inpractice.
4.2 Effects of UHM-CFRP tendonson the structural behaviour of prestressed
concrete beam elements
In 4-point bending, the scatter in the load–deflection
behaviour was found to be much smaller within eachgroup of beams if compared to the analysis of theprestress level and length. In the case of deflection, theoverall beam geometry was dominant for the investigated behaviour. Within the given beam dimensions,the influence of the local material variations, thegeometrical imperfections and the tendon positioning
tolerances were of minor importance and likely tohave resulted in only small deviations within each
respective group of beams.
The tested prestressed concrete beams showed the
expected bi-linear load deflection behaviour; see[43, 44] and Fig. 6a. This bi-linear behaviour could
be described by the flexural stiffness in the respective
section. The flexural stiffness follows the grossmoment of inertia (Ig) in the first branch (M \ Mcr),the cracked moment of inertia (Icr) in the second
branch (M [ Mcr) always in combination with theMaterials and Structures (2021)54:51

elastic modulus (E) of the concrete [43]. For the beams
used in this study, Ig is affected little by the presence of
the CFRP tendons, see Table 3. Consequently, thebehaviour until cracking should be comparable for all
investigated beam configurations, which was indeed
confirmed by the experimental results. When using thetransformed area diagram to calculate Icr, see [45], it isobvious that Icr must be smaller than Ig. In this case thecontribution of the rectangular cross section of theconcrete beam decreases while the contribution of the
CFRP tendons, due to Steiner‘s theorem, increases.
Moreover, the contribution of the CFRP tendons to Icr
is proportional to the modular ratio between tendon
and concrete (n = Ecfrp/Ec), which would logicallyresult in a stiffer deflection behaviour in the cracked
state for UHM-CFRP tendons in comparison to their
UTS-CFRP counterparts. Both expectations wereconfirmed by the experimental results. In fact, therelation between the cracked moments of inertia
Icr_UHM/Icr_UTS was calculated to be 2.9, whichcorresponds well to the ratio of 3.0 found in theexperimental deflection of both beam types.
However, the load bearing capacity of beams
prestressed by UHM-CFRP tendons exceeded thecapacity of beams prestressed by UTS-CFRP tendonsin the experiments. This was unexpected, considering
the results of the tensile tests on the respective
materials (see Table 4). Based on these results, in
combination with the analytical beam deflection
model of Rezazadeh et al. [33], the maximum load
carrying capacity of UTS-CFRP prestressed beams
was calculated to be 11.34 kNm and for UHM-CFRP
beams 11.10 kNm. Both calculations assume a tendon
tensile failure. This could be assumed based on
calculating a critical amount of CFRP. A value smaller
than the critical value would yield towards a tendon‘stensile failure and a value larger to concrete crushing.
For the investigated beams, following [33], the criticalamount of CFRP was calculated to be 2.01% for C1UHM-CFRP prestressed beams and 0.61% for C1UTS-CFRP prestressed beams. The actual amount of
CFRP in the tested beams reached only 0.36%.
Accordingly, the assumption in the model to havetendon tensile failure in both cases was verified.
The employed model by Rezazadeh et al. describedthe experimental load–deflection behaviour withaccuracy (see Fig. 8a). The observed differences in
the load bearing capacity could be based on multiple
reasons. Two of them are discussed in the following.

Materials and Structures (2021)54:51

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51

Table 3 Moments of inertia of cracked and un-cracked beams and the modular ratio between CFRP and concrete
Beam configuration

Ig (mm4)

Icr_exp (mm4)

Icr_cal (mm4)n (–)

C1-UTS-CFRP

2.97 9 107

2.59 9 106

2.65 9 106

4

7

6

C1-UHM-CFRP

3.11 9 107.75 9 107.66 9 106

C4-UHM-CFRP

3.08 9 107

14

n.a

n.a

12

n.a. means not availablen (–) means modular ratio between CFRP and concrete

Table 4 Longitudinal properties of CFRP tendons
Tendon material

Tensile Strength
S11u
(MPa)

SD

Young’s Modulus E11
(GPa)

SD

UHM-CFRP

1562

68

509*

22*

UTS-CFRP

2031

31

145

4

Ultimate tendon stressin beams S11u_beams
(MPa)

SD

1426*

1924

146

1969

1561

34

Design Tensile
Strength S11uD(MPa)*data equivalent to [3]First, the concrete properties themselves could be
affected and hence the critical ratio between the CFRP
tendons and the concrete could have been changed andso the failure mode of the beams. The beam‘s top-fibre
compressive strain during loading was analysed after
the 4 point bending experiments with the help of the
DIC system, Fig. 8b. These results indicate clearly
that the maximum compressive strain in UTS-prestressed beams reached a value of 0.0018 and in UHM-

prestressed beams 0.0012 respectively. Both values
are significantly lower than the suggested theoretical
failure strain of 0.003 given by ACI [44] and aexperimentally tested compressive strains in [46].
Hence it could be concluded that the concrete was far
from crushing and is not expected to be the cause of
the beam‘s early failure.
Secondly, the stress distribution inside the CFRP
tendons in a cracked section of the beam could be

Fig. 8 Influences on a prestressed beam‘s deflection behavior:a Analytical deflection modelling, based on the approach ofRezazadeh et al. [33], in comparison to the experimental results.

b Development of the concrete‘s top fibre strain for C1-UHMCFRP beams and C1-UTS-CFRP-beams during loading

51

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different to the assumed ideal tensile stress condition.
This case would be reasonable if large transverse
stresses were introduced either by the large deflectionsof the UTS-CFRP prestressed beams or due to thedifferent CFRP material properties. Hence, the tendon
could show an unexpected early failure.
To shed light on the stress distribution in the CFRP
tendons, the investigated beams were analysed by asimplified finite element analysis (FEA) in Abaqus 6.14 using a quarter model of a beam‘s cut-out
section with a length of 60 mm, a height of 157.5 mm
and a width of 45 mm. In this model, all materials
exhibited linear elastic behaviour. The CFRP materials were defined by TYPE = ENGINEERING CONSTANTS to consider their anisotropic properties andthe concrete was simply represented by the material
TYPE = ISOTROPIC. The respective properties weretaken from Tables 1, 4 and [6]. Concrete and CFRP
tendons were modelled with elements of TYPE =C3D20R. The beam‘s concrete geometry was cut torepresent the crack until the position of the neutral axis
as observed in the experiments. The uncracked area inthe top section of the beam and the related crosssection of the tendon were restricted by the BOUNDARY CONDITION of TYPE = ZSYMM, whereZ-direction being the normal direction of the ideal
crack surface. The global mesh size was 3.7 mm andthe refined mesh around the crack area was 0.5 mm.
For the LOAD CONDITION, at first the cracked cutout section was prestressed by a PREDEFINEDFIELD of TYPE = STRESS to the same level asfound in the experiments. Afterwards the model was
continuously loaded with the bending moment corresponding to the experimentally found ultimate load
bearing capacity for the respective beam type, seeFig. 6 and Table 5. To account for a possible damageof the bond interface between tendon and concrete, thefirst 3 mm from the crack towards the tendon weremodelled with a SURFACE TO SURFACE CONTACT using a high tangential friction (l = 0.8) of
TYPE = PENALTY and a HARD CONTACT in thenormal direction. The remaining contact of the tendon
was ideally bonded by using TIE to the surrounding
concrete. Time-dependent properties of the concrete,such as creep and shrinkage, were not included in themodel. This simplification was justified by the shorttime frame of the investigated cracked beam situation,shortly before failure. For validation of the model, thecurvature at the ultimate bending moment as well as

Materials and Structures (2021)54:51

the prestress inside the tendon after the prestressrelease was analysed and found to be in good
agreement with the experiments.
At the experimental failure load, the UTS-CFRP
prestressed beams showed a tensile stress in thetendon‘s vertical direction (S22) of [ 50 MPa and the
UHM-CFRP tendons of around 20–25 MPa, Fig. 9.
Both are critical values in the transverse direction, e.g.
in [47] a comparable unidirectional UTS-CFRP
material reached a transverse tensile strength of only
24 MPa. It can be concluded that the low load bearing
capacity of the UTS-CFRP beams was likely due to theunfavourable stress situation in the limit state.
According to the FEA, this was found to be morecritical for the UTS-CFRP beam than for the UHMCFRP beam, see Fig. 9.
4.3 Effects of LCHPC on CFRP prestressed
structural elements
The tested LCHPC recipe (C4) had both Young‘s
modulus and compressive strength very close to thecurrent industry standard C1, see Table 1. Hence, thedeflection behaviour was mainly controlled by the
UHM CFRP prestressing tendons and not by the
LCHPC, see Fig. 6-A. In addition, effects on theprestress loss based on the lower shrinkage and creepof LCHPC (C4) in comparison to HPC (C1) were
expected, see [6], but could not be identified in theexperimentally investigated prestressed concretebeams. This was mostly caused by the high scatter inthe experimental results observed in the C4-UHMCFRP beam specimens, see Fig. 4.
Recently, in [6], the effects of shrinkage, creep andelastic shortening on prestress losses in LCHPC
prestressed elements were estimated based on an
idealized and simplified FEA. The geometry of thestudied elements was based on a shrinkage and creep
test specimen to reduce geometrical effects of thespecimen on modelling shrinkage and creep and hence
to allow a comparison between modelling and therespective experiments. The beams used in this study
did not have the same dimensions as the specimens
examined in [6]. The cross-sectional area of theemployed beams was about 2 times larger and they
were about 8 times longer than the creep and shrinkage
specimens in [6]. In a first step, the Abaqus 6.14 FEA
model, developed in [6], was adapted in geometry andapplied to the beam specimens tested in this study. The

Materials and Structures (2021)54:51

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51

Table 5 Summarized results of beam tests
Beam
configuration

Cracking
load (kN)Midspandeflection
cracking dc(mm)

Stiffness
unracked(kN/mm)Ultimateload (kN)

Ultimate
midspandeflection du(mm)

Ultimate
tendonstrain (%)

Stiffness
cracked(kN/mm)Failure
mode
(–)

du/
dc
(–)

C1-UTS B1

12.43

6.37

1.95

21.65

37.42

1.06

0.30

TF

5.87

C1-UTS B2

12.29

5.99

2.05

21.85

38.28

n.a

0.30

TF

6.39

C1-UHM B1

10.76

3.56

3.02

26.59

21.44

0.39

0.89

TF

6.02

C1-UHM B2

11.13

3.63

3.07

26.42

20.79

0.35

0.89

TF

5.73

C1-UHM B3

11.16

3.56

3.13

25.91

19.58

n.a

0.92

TF

5.50

C4-UHM B1
C4-UHM B2

11.84
11.66

3.81
2.87

3.11
4.06

27.68
25.98

20.64
16.88

0.40
n.a

0.94
1.02

TF
TF

5.42
5.88

C4-UHM B3

9.04

3.20

2.83

25.43

18.91

n.a

1.04

TF

5.91

TF means tendon failuren.a. means not available

Fig. 9 FEA results showing the vertical stress component (S22) in the CFRP tendons at ultimate load condition. a For a C1-UHMCFRP beam. b For a C1-UTS-CFRP prestressed beam

initial prestress of 800 MPa was applied to the tendons
with a PREDEFINED FIELD of TYPE = STRESS,
shrinkage of concrete was modelled by a thermal
expansion coefficient analogy and creep by thefunction *VISCOELSTIC. Further details about the
FEA can be found in [6]. Prestress losses in the FEA
were determined by comparing the remaining prestress in the tendons at a certain time with the initialprestress applied to the tendons. For validation, theresults of prestress losses after 28 days were comparedto the measured prestress losses, see Fig. 10a. The
contribution of the elastic shortening, shrinkage andcreep on the prestress loss were analysed based on the
FEA, Fig. 10b. As in [6], it was found that all threecontributors deliver noteworthy amounts of prestressloss to the investigated beam specimens. Based on the

FEA, the prestress loss in LCHPC beams was calculated to be only around 3% less in comparison to their
HPC beam counterparts, see Fig. 10a and b. Considering the large scatter in the experimental prestressloss measurements, in particular for C4-UHM-CFRP
beams with results based on only two beams, thiseffect could neither be disproved nor confirmed by theexperiments.
As a result, the theoretical benefit of a higher
remaining prestress in the structure due to the lower
creep and shrinkage was not significant, not even in
the case when the tensile-wise very sensitive UHMCFRP tendons were used for prestressing. However,
the experiments described in this paper confirm that
the LCHPC, with their clinker reduction of up to 70%,
could be used as a replacement for the current HPC in

51

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Materials and Structures (2021)54:51

Fig. 10 Prestress loss assessment. a Comparison of experimental results with FEA predictions. b Breakdown of prestress losses into thethree main contributions of shrinkage, creep and elastic shortening based on the FEA

CFRP prestressed structural elements without compromising their short term mechanical performance.
These results are considered to be valid for the specificcase investigated in this study. The effects of creep andshrinkage might be different in structures of otherdimensions, which would need further investigations
that are beyond the scope of this work.

5 Conclusions
Based on monitoring the prestress inside beam specimens up to 28 days after casting, full scale 4-point
bending beam tests and additional upscaling of a small
scale FEA model to structural level in Abaqus 6.14,
the following conclusions can be drawn:• The prestress transfer length is affected by tendon
stiffness. Here, the shear stress transfer between
the prestressing tendon and the concrete is dependent mainly on the stiffness of the materials in the
direction of the applied prestress. Hence, in thesame concrete matrix, the stiffer tendon has thelonger transfer length.
• Beams prestressed with stiffer tendons experiencelarger prestress losses. Elastic shortening, shrinkage and creep all contribute to a shortening of theconcrete beam in the prestress direction. A stiffer
prestressing tendon, embedded and bonded to theconcrete, will always experience more prestressloss due to the same amount of shortening.

• In the cracked state UHM-CFRP prestressed
beams showed less deflection, narrower cracks
and an increased load bearing capacity in comparison to their UTS-CFRP prestressed opponents.
When using UHM-CFRP, the beams could be used
more efficiently in terms of their structural
performance.
• LCHPC, with much lower clinker content than
traditional HPC, is equally suited for CFRPprestressing applications. In particular, the flexural
behaviour of prestressed beams was found to beequivalent to that of the reference HPC.
The results of this work show very good potential
for improving the current CFRP prestressing technology at the structural level by using UHM tendons.
While LCHPC has lower creep and shrinkage, in theparticular structure investigated in this study no
benefits in term of reducing prestress losses wasobserved, possibly due to the high scatter of thesemeasurements. However, these results still focus
mainly on the short-term behaviour of a prestressed
beam. Critically, also data about the structural
behaviour and its integrity over a longer period oftime (several years) in harsh environmental conditions
would be needed. This would possibly help industrial
stakeholders to gain confidence in these new materials
and finally bring this technology into construction
practice.
Acknowledgements This research project was part of thejoint project ‘‘Low energy concrete’’ within the framework ofMaterials and Structures (2021)54:51
the National Research Programme ’’Energy Turnaround‘‘ (NRP70) of the Swiss National Science Foundation (SNSF). Further
information on the National Research Programme can be foundat http://www.nrp70.ch/en/projects/building-settlement/lowenergy-concrete. We are grateful to Daniel Käppeli, Sebastiano Valvo, Marcel Käppeli, and Daniel Völki for their strong
support during sample production and testing.
Funding Open Access funding provided by Lib4RI – Libraryfor the Research Institutes within the ETH Domain: Eawag,
Empa, PSI & WSL.
Open Access This article is licensed under a Creative
Commons Attribution 4.0 International License, which permitsuse, sharing, adaptation, distribution and reproduction in any
medium or format, as long as you give appropriate credit to theoriginal author(s) and the source, provide a link to the Creative
Commons licence, and indicate if changes were made. The
images or other third party material in this article are included in
the article’s Creative Commons licence, unless indicatedotherwise in a credit line to the material. If material is notincluded in the article’s Creative Commons licence and yourintended use is not permitted by statutory regulation or exceeds
the permitted use, you will need to obtain permission directlyfrom the copyright holder. To view a copy of this licence, visit
http://creativecommons.org/licenses/by/4.0/.

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