CORROSION INDUCED FAILURES OF PRESTRESSING STEEL ...

Corrosion induced failures of prestressing steel 

CORROSION INDUCED FAILURES OF PRESTRESSING STEEL 

KORROSIONSBEDINGTE VERSAGENSMECHANISMEN BEI 

SPANNSTAHL 

RUPTURES D'ARMATURE DE PRECONTRAINTE INDUITES PAR 

CORROSION 

Ulf Nürnberger 

SUMMARY 

Rarely in prestressed concrete structures occurring fractures of prestressing 

steel in prestressed concrete structure can, as a rule, be attributed to corrosion 

induced influences. The mechanism of these failures often is not well understood. 

In this connection it is difficult to establish the necessary recommendation 

not only for design and execution but also for building materials and prestressing 

systems in order to avoid future problems. This paper gives a survey about 

corrosion induced failure mechanisms of prestressing steels with a particular 

emphasis on post-tensioning tendons. 

Depending on the prevailing corrosion situation and the load conditions as 

well as the prestressing steel properties the following possibilities of fracture 

must be distinguished: 

• Brittle fracture due to exceeding the residual load capacity. Brittle fracture is 

particularly promoted by local corrosion attack and hydrogen embrittlement. 

• Fracture as a result of hydrogen induced stress-corrosion cracking. 

• Fracture as a result of fatigue and corrosion influences, distinguishing between 

corrosion fatigue cracking and fretting corrosion/fretting fatigue. 

ZUSAMMENFASSUNG 

Die gelegentlich an den im Spannbetonbau verwendeten Spannstählen auftretenden 

Brüche sind im Regelfall auf korrosionsbedingte Einflüsse zurückzuführen. 

Die Versagensmechanismen werden häufig nicht ausreichend verstanden. 

Deshalb ist es schwierig, die notwendigen Empfehlungen nicht nur für Planung 

und Ausführung sondern auch für die Auswahl der Baustoffe und Vorspannsysteme 

zu geben, um zukünftige Probleme auszuschließen. Der Beitrag 

9 

Otto-Graf-Journal Vol. 13, 2002

U. NÜRNBERGER 

stellt in einem Überblick die korrosionsbedingten Versagensmechanismen von 

Spannstählen, mit Schwerpunkt der Probleme bei nachträglich vorgespannten 

Zuggliedern, dar. 

In Abhängigkeit sowohl von der vorherrschenden Korrosionssituation und 

den Belastungsverhältnissen als auch den Spannstahleigenschaften müssen die 

folgenden Brucharten unterschieden werden: 

• Sprödbruch durch Überschreiten der Resttragfähigkeit. Das Auftreten eines 

Sprödbruches wird unterstützt durch einen lokalen Korrosionsangriff und eine 

Wasserstoffversprödung. 

• Bruch infolge wasserstoffinduzierter Spannungsrisskorrosion. 

• Brüche als Folge von Ermüdung und Korrosionseinflüssen. Hierbei ist zu 

unterscheiden zwischen Schwingungsrisskorrosion und Reibkorrosion/Reibermüdung. 

RESUME 

Les ruptures occasionnelles des armatures de précontrainte peuvent en général 

être attribués à l'influence de la corrosion. Le mécanisme de ces ruptures 

n'est souvent pas bien compris. Il est par conséquent difficile d'émettre des recommandations, 

non seulement pour la conception et l'exécution, mais également 

pour le choix des matériaux et des systèmes de précontrainte. Cet article 

donne un aperçu sur les mécanismes de ruptures induites par corrosion des 

armatures de précontrainte, en particulier sur les armatures précontraintes par 

post-tension. 

En fonction des conditions corrosives de l'environnement, des conditions 

de chargement et des propriétés de l'armature précontrainte, on distingue les types 

de rupture suivants: 

• rupture fragile due au dépassement de la capacité résiduelle de charge. La 

rupture fragile est favorisée par la corrosion locale et la fragilisation par hydrogène. 

• rupture par corrosion sous contrainte induite par l'hydrogène. 

• rupture en raison des influences combinées de fatigue et corrosion. On distingue 

la fatigue sous corrosion et la corrosion par friction/fatigue par friction. 

KEYWORDS: prestressed concrete, corrosion, failures, steel 

10

1. INTRODUCTION 


Most of the prestressed concrete structures built in the last 50 years in accordance 

with the rules for good design, detailing and practice of execution have 

demonstrated an excellent durability [1]. Analyses of occasional problems confirm 

that instances of serious failures are rare considering the volume of 

prestressing steels that has been in use worldwide. 

Major issues which strongly influence the level of durability actually 

achieved are insufficient design (poor construction), incorrect execution of 

planned design (poor workmanship), unsuitable mineral building materials, unsuitable 

post-tensioning system components, including the prestressing steel [1- 

3]. Insufficient design and incorrect work execution will mean that the necessary 

corrosion control is not guaranteed from the beginning in all areas or that as a 

result of natural influences (i. e. carbonation, chloride ingress) it will get lost 

soon within the time frame of the originally anticipated life time. Unsuitable materials 

or inappropriate substances in materials will further corrosion and/or 

stress corrosion cracking. Sensitive prestressing steels cannot withstand even 

inevitable building-site influences or will fail while in use. 

Most corrosion defects are caused by water which seeps through zones of 

porous concrete and vulnerable areas such as leaking seals, joints, anchorages or 

cracks, and which flows through the network of ducts which have been grouted 

to a greater or lesser extent. The major threat is corrosion due to chlorides. The 

source of chlorides can be either de-icing salts or seawater. 

Rarely occurring fractures of prestressing steel and failures of prestressed 

concrete structure can, as a rule, be attributed to corrosion induced cracking. The 

mechanism of these failures often is not well understood. In this connection it is 

difficult to establish the necessary recommendation not only for design and execution 

but also for building materials and prestressing systems in order to avoid 

future problems. 

This paper gives a survey about corrosion induced failure mechanisms of 

prestressing steels with a particular emphasis on post-tensioning tendons. 

11 



2. FRACTURE MECHANISMS OF PRESTRESSING STEEL 

The types of corrosion occurring at times as well as their specific manifestation 

must be regarded as an essential influencing factor on the behaviour of the 

prestressing steels under unforeseen or inappropriate service conditions. The 

exclusive determination that corrosion was involved is not enough for a critical 

case study and for future damage prevention. 


well as the prestressing steel properties the following possibilities of fracturing 


• Brittle fracture due to exceeding the residual load capacity. Brittle fracture is 

particularly promoted by: 

− local corrosion attack (pitting and wide pitting corrosion), 

− hydrogen embrittlement. 

• Fracture as a result of stress corrosion cracking, where we distinguish between 

− anodic stress corrosion cracking and 

− hydrogen induced stress-corrosion cracking. 


− corrosion fatigue cracking and 

− fretting corrosion/fretting fatigue. 

In the following such events will be described in more detail, also with regard 

to prestressed concrete construction. 

2.1 Brittle fracture 

Brittle fracture may occur in high-strength steels after swift tensile stress. 

This is the case in prestressing steels when there is a fracture under loads until 

reaching the permissible pre-strain as a result of these influences: 

− stress concentration in local notches (e. g. wide corrosion pit), 

− high stressing speed and low temperature, 

− an embrittlement of the steel structure after hydrogen adsorption 

(hydrogen embrittlement). 

12

Influence of corrosion 


Mainly uniform general corrosion (e. g. after a prolonged weathering on a 

building site) does not have any major impact on the load bearing capacity. Not 

until, due to corrosion, an underrun of the required residual cross section has 

taken place than a prestressing steel fracture may occur after exceeding the residual 

load bearing capacity. Such events may happen once prestressing steels in 

ungrouted tendon ducts are exposed over a long period of time to water and 

oxygen via untight anchorages or construction joints. 

If, however, the prestressing steel incurs a local corrosion attack in the 

form of pitting or wide pitting corrosion, the load bearing capacity may get lost 

at an early stages due to brittle fracture. The following effects are capable of 

triggering such attacks in prestressing steel: 

The presence of aggressive water in the not yet injected ducts of post tensioning 

tendons which result from bleeding of the concrete during the erection 

of the construction. Already in the not grouted and not prestressed condition the 

steel may suffer from strong pitting or wide pitting corrosion and the load bearing 

capacity can be reduced considerably. 

Bleeding is a separation of fresh concrete, where the solid content sinks 

down and the displaced water rises or penetrates in the inner hollows. In the 

bleeding water significantly high contents of sulphates and increased quantities 

of chlorides may be accumulated (Table 1) by leaching of the construction materials 

cement, aggregates and water. The high amounts of potassium-sulphates 

result from the gypsum in the cement. The watery phase of fresh concrete penetrates 

into the ducts through the anchorages, couplings and defects in the sheet 

and accumulates at the deepest points. Because of an access of air the alkaline 

water carbonates quickly. As early as in the non-grouted and non-prestressed 

condition the steel can suffer from strong pitting. Bleeding water attack may 

within a few weeks lead to pitting depths of up to 1 mm. 

Table 1: Analysis of bleeding water 

sulphate 1.90 - 5.20 g/l 

chloride 0.13 - 0.18 g/l 

calcium 0.06 - 0.09 g/l 

sodium 0.18 - 0.37 g/l 

potassium 3.60 - 7.30 g/l 

pH-value 10 - 13 

13 



The access of chloride containing waters, e.g. above untight anchorages or 

joints, in a non-grouted tendon duct may lead to damaging local corrosion attack 

in prestressing steel during the life time and after years of use. Comparable attacks 

must be expected once chloride salts penetrate to the tendon through a 

concrete cover of inferior thickness and impermeability. 

The performance characteristics of corroded prestressing steels can be determined 

in tensile, fatigue and stress corrosion tests (Fig. 1). Such tests to establish 

the residual load bearing capacity will, for instance, be carried out while inspecting 

older buildings, after damaged prestressing steel samples had been 

drawn. This might help to gain the knowledge for necessary repair. 

High strength prestressing steels show a far more sensitive reaction to corrosion 

attack than reinforcing steels, and this increasingly in the sequence tensile 

test - fatigue test – stress corrosion test [4]. In case of uneven local corrosion a 

corrosion depth of 0.6 mm may suffice for breaking a cold deformed wire under 

tension of 70 % of the specified tendon strength of about 1800 N/mm 2 (Fig. 1, 

tensile test). 

At pitting depth of above 0.2 mm cold drawn wires may show fatigue limits 

(fatigue limits for stress cycles of N = 2 · 10 6 ) of 100 N/mm 2 and less (Fig. 1, 

fatigue test). Like-new smooth surfaced steels normally show a fatigue limit of 

more than 400 N/mm 2 . 

In all the performance characteristics of prestressing steels local corrosion 

attack has the most detrimental effect on the behaviour to hydrogen induced corrosion 

cracking. In a test developed by FIP the prestressing steel is immersed 

under tension into an ammonium thiocyanate solution. A minimum and average 

time of exposure before failure is specified. For cold drawn wire and strand 

these values are in the order of 1.5, respectively 5 hours. In this example these 

life times are underrun at corrosion depths of > 0.2 mm (Fig. 1, stress corrosion 

test). 

14

tensile strength R m in N/mm² 

2000 

1600 

1200 

800 

400 

lifetime in h 

tensile test 

0 

0 

0,0 0,2 0,4 0,6 0,8 1,0 

10 

8 

6 

4 

2 

0 

A 10 

R m 

depth of uneven local corrosion 

in mm 

stress corrosion test 

FIP - test: 

20% NH 4SCN 

50° C 

0,0 0,2 0,4 0,6 0,8 1,0 


in mm 

10 

8 

6 

4 

2 

elongation at fracture A 10 in % 

fatigue loading in N/mm² 

for N=2*10 6 

Effect of hydrogen (hydrogen embrittlement) 


500 

400 

300 

200 

100 

0 

fatigue test 

0,0 0,2 0,4 0,6 0,8 1,0 


in mm 

Fig. 1: Properties of cold-deformed 

prestressing steel wires St 1570/1770, 

dS = 5 mm, in relation to depth of uneven 

local corrosion [Nürnberger] 

(scattering of 90% of the test results) 

In a specific corrosion situation prestressing steel corrosion may release 

hydrogen which is then absorbed by the prestressing steel, which, if prestressed 

at the same time, will allow hydrogen induced stress corrosion cracking with 

crack initiation and crack propagation (chapter 2.2). Also if the prestressing steel 

is free of any tensile stresses (not prestressed), hydrogen can be absorbed in the 

event of corrosion. The steel will not crack, but depending on the quantity of 

hydrogen absorbed and the specific hydrogen sensitivity the prestressing steel 

may become brittle. This may have an adverse effect on the mechanical characteristics 

[5], more so on the deformation properties than on the tensile strength. 

15 



reduction in area Z in % 

50 

40 

30 

20 

10 

Z 

R m 

0 

-100 

0 10 1 20 2 30 5 40 10 50 20 50 60 

time of immersion in testing medium 

according FIP (20% NH 4 SCN, 50° C) in h 

2000 

1825 1900 

1650 1800 

1475 1700 

1300 1600 

1125 1500 

950 

775 

600 

425 

250 

Fig. 2 Tensile strength (Rm) and reduction in area (Z) of cold deformed prestressing steel 

wires (4 steel melts) after charging with hydrogen [5] 

Prestressing steel fractures as a result of corrosion-caused hydrogen embrittlement 

may occur, for instance, when prestressing to a high stress level or 

shortly after the prestressing, after the steel had been absorbing high quantities 

of hydrogen in an enduring unfavourable corrosion situation. If properly and 

swiftly processed, such damages, indeed, should not occur. 

2.2 Fractures because of stress-corrosion cracking 

Stress-corrosion cracking is understood to mean crack formation and crack 

propagation in a material under the effect of mechanical tensile stresses and of 

an aqueous corrosion medium. 

16 

75 

tensile strength R m in N/mm²

Anodic stress-corrosion cracking 


In the presence of nitrate-containing non-alkaline electrolytes (pH-value 

< 9) unalloyed and low-alloy steels may suffer an anodic stress-corrosion cracking. 

Crack formation and crack propagation are due to a selective metal dissolution 

(e. g. along grain boundaries of the steel structure) with a simultaneous effect 

of high mechanical tensile stresses [6] on condition that there is special tendency 

of the steels to passivate in nitrate-containing aqueous solutions. 

In the prestressed concrete construction the media-related pre-conditions, 

e.g. in the fertilizer storage and in stable ceilings, can be assumed as a fact. In 

stables brickwork, salpetre Ca (NO3)2 may be formed by urea. In the presence of 

moisture the nitrates may diffuse into the concrete and may cause stresscorrosion 

cracking in the case of pretensioned concrete components affecting the 

tension wires if the concrete cover is carbonated due to an inferior quality of the 

concrete [6]. 

A specific nitrate sensitivity of the steels is always a pre-condition for an 

anodic stress-corrosion cracking. Low-carbon concrete steels are very susceptible 

to nitrate induced stress-corrosion cracking. The prestressing steels currently 

in use, however, are highly resistant to this type of corrosion. 

Hydrogen induced stress corrosion cracking [6,7] 

Fractures of prestressing steel as a rule can be referred to hydrogen induced 

stress corrosion cracking (H-SCC). It may happen during the erection of the 

construction or during later use. The following conditions are necessary: 

• a sensitive material or state, 

• a sufficient tension load, 

• at least a slight corrosion attack. 

The risk of fractures due to hydrogen induced stress corrosion cracking 

therefore results from the joint action of very prestressing steel properties and 

environmental parameters. What is needed is the presence of hydrogen which 

comes into being under certain corrosion conditions in neutral and particularly 

in acid aqueous media through the cathodic partial reaction of the corrosion. 

17 



A reduction of oxygen access may support evolution of adsorbable atomic 

hydrogen (then reaction 6 is hindered). Therefore at the surface of corroding 

steel the amount of adsorbable hydrogen atoms rises 

• with increasing hydrogen concentration (reaction 3 or 4 is accelerated), 

• in the presence of so-called promotors (reactions 5 is hindered), 

• in an electrolyte impoverished in oxygen (reaction 6 is hindered). 

From the practical point of view one can say that hydrogen assisted damages 

are only possible 

• in acid media or if the steel surface is polarized to low potentials (e. g. if 

the prestressing steel has contact with zinc or galvanized steel), 

• in the presence of promotors such as sulphides, thiocyanate or compounds 

of arsenic or selenium, 

• and under crevice conditions, because the electrolyte in the crevice is poor 

in oxygen. 

Fig. 3: Pitting induced stress corrosion cracking 

In concrete structures the attacking medium is mostly alkaline and acid 

media are limited to exceptions. Nevertheless, in natural environments the pitting 

induced H-SCC can take place (Fig. 3). Pitting induced H-SCC means crack 

initiation within a corrosion pit. In the corrosion pits the pH-value falls down 

because of hydrolysis of the Fe 2+ -ions. Pitting or spots of local corrosion can be 

explained by differential aeration or concentration cells. Especially effective is 

19 



the attack of condensation water or salt enriched aqueous solution (bleed water, 

chapter 2.1), when erecting the constructions. 

In prestressed construction chloride contamination supports a local corrosion 

attack. In the case of sensitive prestressing steel all but minimal contents of 

hydrogen can lead to irreversible damages. Then a minimal local corrosion attack 

without visible corrosion products on the steel surface may lead to steel 

fracture. 

In prestressed concrete structures all types of uneven local corrosion should 

be prevented to exclude failures because of hydrogen assisted cracking. 

The preconditions for "classical" stress-corrosion cracking are most readily 

to be found in prestressed concrete construction, i. e. crack formation and 

propagation under purely static stress. By prestressing the stress amplitudes of 

the structure caused e. g. by wind and traffic are kept low. Nevertheless, the occurrence 

of pulsating loads or service-related strain changes of the steels will 

raise the crack corrosion risk since it will favour hydrogen induced "nonclassical" 

stress-corrosion cracking [6]. Plastic flow in steel favours an absorption 

of atomic hydrogen. 

stress amplitude 2 σA in N/mm 2 

200 

175 

150 

125 

100 

75 

50 

25 

0 

10 5 

air 

aqueous solution 

2 4 6 8 

10 6 

agent: 

2 4 6 8 

number of stress reversal 

5 g/ 

l K 2SO4 

with 

0, 

5 g/ 

l KCl without 

RT 

without failure 

10 7 

1g/l NH 4SCN 

2 4 6 8 

5 11 22 55 111 222 555 1111 2222 5555 

lifetime in hours 

Fig. 4: SCC-behaviour of prestressing steel St 1420/1570 (German standard) ∅ 12.2 mm 

without and with dynamic stress of low amplitude 

20 

10 8


Fig. 4 [8] compares the behaviour of a quenched and tempered prestressing 

steel (from a case of damage) sensitive to hydrogen in a stress-corrosion cracking 

test with and without superimposed fatigue loading of low amplitude (30 – 

80 N/mm 2 ). The aqueous test solution contains 5 g/l SO4 2- ,0.5 g/l Cl¯ without 

and alternatively with 1 g/l SCN¯ as a promotor for a hydrogen absorption. The 

stress-corrosion cracking test under static stress was realized at 80 % of the tensile 

strength. This stress corresponds to the constant maximum stress in the tensile 

fatigues test. Fig 4 represents the stress cycle number as a function of the 

amplitude, in the course of which also the life time, calculated over the frequency 

(f = 5s -1 ), is applied. The stress corrosion test results without superimposed 

fatigue loading are applied at a range of stress of 0 N/mm 2 . The hydrogen 

insensitive steel failed in the "static" test within a test period of 5000 hours in 

the promotor-containing solution but did not fail in the promotor-free solution. If 

a fatigue test of low amplitude is superimposed, the lifetime in the promotorcontaining 

solution will more and more decrease with rising amplitude. In the 

wave stress it is striking that fractures also occur on steels in the promotor-free 

solution. 

It was found that in cold deformed prestressing steels the influence of a superimposed 

fatigue loading on the hydrogen induced stress-corrosion cracking is 

revealing itself weaker. These tests lead to the conclusion that already fatigue 

loadings of low amplitude or elongations caused by changes in utilization tend 

to significantly jeopardize the susceptibility of prestressing steels to stresscorrosion 

cracking. 

2.3 Fractures because of fatigue and corrosion 

Prestressing steels can only be subject to a noticeable steel stress in dynamically 

strained reinforced concrete structures if there is concrete in a cracked 

state. The stress amplitudes of prestressing steel due to acting high dynamic 

loads ( e. g. a high traffic load of a bridge) may then amount to > 200 N/mm 2 in 

the crack region. In the uncracked state the steels will show ranges of stress of 

clearly less than 100 N/mm 2 . 

Cracks in concrete may occur in partially prestressed structures. Since such 

cracks tend to open and to close in a superimposed fatigue stress the following 

facts must be considered: 

21 



Corrosion fatigue cracking 

If corrosion promoting aqueous media penetrate through the concrete crack 

to the dynamically stressed tendon, corrosion fatigue cracking is possible although 

this type of corrosion has not been observed in prestressing steel construction 

so far. Corrosion fatigue cracking [6] manifests itself in that a metallic 

material under dynamic stress in a reactive corrosion medium (water, salt solution) 

will show a much more unfavourable fatigue behaviour than under fatigue 

loading in air. This can be explained by characteristic interactions of metal 

physical and corrosive processes which favour initial precrack formation and 

propagation. As opposed to the stress-corrosion cracking the corrosion fatigue 

cracking does not require a specifically acting corrosion medium. 

In case of post-tensioning tendons the duct made of thin steel sheets does 

not offer a lasting corrosion protection and may even suffer fatigue fractures under 

dynamic stress [9]. 

A decrease of the fatigue limit by corrosion is the more distinct the higher 

the strength of the steel and the more aggressive an attacking medium are. 

Hence the high strength prestressing steels, when e. g. simultaneously attacked 

by an aqueous chloride-containing medium, may show a very unfavourable fatigue 

behaviour. 

In traffic carrying bridge structures only the low-frequent stresses lead to 

high stress amplitudes. This results in additional unfavourable conditions with 

regard to corrosion fatigue cracking: with a falling frequency the influence corrosion 

will increase and the fatigue limit will consequently drop. 

For a cold drawn prestressing steel wire Fig. 5 shows a decrease of the corrosion 

fatigue limit in the sequence air-water-chloride solution. For frequencies 

of 0.5s -1 the fatigue limit for stress cycles of 10 7 is below 100 N/mm 2 . 

The problem of corrosion fatigue cracking of cracked components can be 

remedied by sufficient concrete cover and limiting the crack width. This is the 

way of keeping pollutants away from the prestressing steel surface. 

Fretting corrosion / fretting fatigue 

In the vicinity of concrete cracks due to fatigue loading displacements between 

the tendon and the injection mortar or the steel duct respectively will occur 

in a cracked component. In bended tendons a high radial pressure acts at the 

22


same time on the fretting prestressing steel surface. If air or oxygen advance to 

the fretting location through the concrete crack a fretting corrosion is favoured 

[6,10]. Fretting corrosion is described as damaging a metal surface similar to 

wear as a result of oscillating friction under radial pressure with a partner. In the 

presence of oxygen oxidation of the reactive surface will take place. 

In fatigue loaded steels and under fretting corrosion stress at the same time 

the fatigue behaviour is under a very unfavourable influence due to fretting fatigue 

[10]. This is attributable to structural disintegration and the occurrence of 

additional tensile strengths in the fretting area. In concrete embedded tendons, 

subjected to a relative movement and a radial pressure in the concrete crack between 

prestressing steel and duct or injection mortar respectively, tolerable fatigue 

limits of about 150 N/mm 2 for cycles to fracture of 2 x 10 6 were found 

[9,11]. 

In prestressed concrete constructions also the anchorages of the tendons, 

due to fretting corrosion influences, show a fatigue limit which is reduced compared 

with the free length [12]. Under dynamic stress of the anchored tendon the 

fatigue limit, depending on the type of anchorage, is reduced to values between 

80 and 150 N/mm 2 . For this reason, anchorages will always be positioned in areas 

of least stress changes. In the fatigue experiment the prestressing steels always 

fracture in the force transmitting area, i. e. at the beginning of the anchorage. 

Here, the fatigue limit is reduced due to the presence of shifting between 

the prestressing steel and the anchor body and the high radial pressures at the 

same time. 

In prestressed concrete bridges, however, particularly the coupling joints 

proved to be problematic. If such joints crack as a result of imposed stresses 

(e.g. due to non uniform sun heating and low amount of reinforcement which 

crosses the coupling joint) the tendon couplings will suffer major stress fatigue 

cycles from the traffic load which also led to prestressing steel fractures owing 

to the stress-sensitive couplings [2,11]. 

23 



Fig. 5: Fatigue behaviour under pulsating tensile stresses of cold drawn prestressing steel 

wires (Rm ≈ 1750 N/mm 2 ) in air and corrosion- promoting aqueous solutions (Nürnberger) 

3. CONCLUSION 


well as the prestressing steel properties the following possibilities of fracturing 


• Brittle fracture due to exceeding the residual load capacity. Brittle fracture 

is particularly promoted by local corrosion attack and hydrogen embrittlement. 

• Fracture as a result of hydrogen induced stress-corrosion cracking. 


corrosion fatigue cracking and fretting corrosion/fretting fatigue. 

REFERENCES 

[1] Durability of post-tensioning tendons. Proceedings of workshop held in 

Gent University on 15 - 16 November 2001. fib technical report, bulletin 

15 

[2] Nürnberger, U.: Analyse und Auswertung von Schadensfällen bei Spannstählen. 

Forschung, Straßenbau und Straßenverkehrstechnik 308 (1980) 1 – 

195 

24


[3] Nürnberger, U.: Influence of material and processing on stress corrosion 

cracking of prestressing steel (case studies). Publ. of fib commission 9.5, to 

be published 

[4] Neubert B., Nürnberger, U.: Erkennen von Spannverfahrensschädigung – 

Untersuchung der statischen und dynamischen Kenngrößen in Abhängigkeit 

von Rostgrad. Bericht II.6-13675 der FMPA Baden-Württemberg, Otto-Graf-Institut, 

Stuttgart 31.01.1983 

[5] Nürnberger, U., Beul, W.: Entwicklung einfacher und reproduzierbarer 

Prüfverfahren für die Empfindlichkeit von Spannstählen gegenüber Spannungsrisskorrosion. 

Bericht 34-14071 der FMPA Baden-Württemberg, Otto-Graf-Institut, 

Stuttgart 01.03.1996 

[6] Nürnberger, U.: Korrosion und Korrosionsschutz im Bauwesen. Bauverlag, 

Wiesbaden 1995 

[7] Grimme, D., Isecke, B., Nürnberger, U., Riecke, E. M., Uhlig, G.: Spannungsrisskorrosion 

in Spannbetonbauwerken. Verlag Stahleisen mbH, Düsseldorf 

1983 

[8] Nürnberger, U., Beul, W.: Wasserstoffinduzierte Spannungsrisskorrosion 

von zugschwellbeanspruchten Spannstählen, S. 302 – 309; in "Bewehrte 

Betonbauteile unter Betriebsbedingungen". Wiley-VCH Verlag 

[9] Cordes, H.: Dauerhaftigkeit von Spanngliedern unter zyklischen Beanspruchungen. 

Sachstandsbericht. Schriftenreihe Deutscher Ausschuß für Stahlbeton 

370 (1986) 

[10] Patzak, M.: Die Bedeutung der Reibkorrosion für nichtruhende Verankerungen 

und Verbindungen metallischer Bauteile des konstruktiven Ingenieurbaus. 

Dissertation Universität Stuttgart, 1979 

[11] König, G., Maurer, R., Zichner, T.: Spannbeton-Bewährung im Brückenbau. 

Springer Verlag Berlin-Heidelberg-New York-London-Paris-Tokyo, 

1986 

[12] Rehm, G., Nürnberger, U., Patzak, M.: Keil- und Klemmverankerungen für 

dynamisch beanspruchte Zugglieder aus hochfesten Stählen. Bauingenieur 

52 (1977) 287 – 298 

25 



26

Load bearing behaviour of fastenings with concrete screws 

LOAD BEARING BEHAVIOUR OF FASTENINGS WITH CONCRETE 

SCREWS 

TRAGVERHALTEN VON BEFESTIGUNGEN MIT SCHRAUBDÜBELN 

COMPORTEMENT SOUS CHARGE DES ANCRAGES AVEC VIS 

D'ANCRAGE 

Jürgen H. R. Küenzlen and Rolf Eligehausen 

SUMMARY 

Concrete screws are a relatively new fastening system. Their main 

advantage compared to traditional post-installed fastening systems is a quick and 

easy installation. A hole is drilled into the concrete and threads are cut in the 

concrete by the screw as it is installed. 

Concrete screws transfer tensile loads into the base material by mechanical 

interlock of the threads. Due to their load-bearing mechanism, concrete screws 

with a technical approval of the DIBt can be used for fastenings in cracked and 

non-cracked concrete. 

The typical failure mechanism for concrete screws is concrete-cone failure. 

With increasing embedment depth the ratio of the depth of the concrete failure 

cone to the embedment depth decreases. The failure load of concrete screws 

with continuous threads along the entire embedment depth increases 

proportionally to hef 1,5 (hef = effective embedment depth), but it is about 20 % 

smaller than the failure load of expansion and undercut anchors with the same 

embedment depth. 

In order for concrete screws to function properly, the threads cut into the 

wall of the drilled hole must not be damaged during the installation. This 

requirement is achieved by using the embedment depth defined in the Technical 

Approvals. 

27 


J. H. R. KÜENZLEN, R. ELIGEHAUSEN 


Schraubdübel sind ein relativ neues Befestigungssystem. Ihr großer Vorteil 

liegt in der einfachen und schnellen Montage. Es wird ein Loch in den Beton 

gebohrt, in das der Schraubdübel beim Setzen ein Gewinde schneidet. 

Schraubdübel werden in Durchsteckmontage gesetzt. 

Schraubdübel übertragen eine angreifende Zuglast über mechanische 

Verzahnung der Gewindeflanken, die in die Bohrlochwand einschneiden, in den 

Untergrund. Aufgrund ihres Tragmechanismus sind bauaufsichtlich zugelassene 

Schraubdübel für Befestigungen im ungerissenen und gerissenen Beton 

geeignet. 

Das Versagen erfolgt durch Betonausbruch, wobei mit zunehmender 

Verankerungstiefe das Verhältnis von Tiefe des Ausbruchkegels zu 

Verankerungstiefe abnimmt. Die Bruchlast steigt bei Schraubdübeln mit einem 

über die gesamte Verankerungstiefe durchgehende Gewinde proportional zu 

hef 1,5 an (hef = Verankerungstiefe), jedoch ist sie unter sonst gleichen 

Verhältnissen ca. 20 % niedriger als die Betonausbruchlast von Spreiz- und 

Hinterschnittdübeln. 

Damit Schraubdübel ordnungsgemäß funktionieren, dürfen die in den 

Beton geschnittenen Gewindegänge nicht während der Montage beschädigt 

werden. Diese Bedingung wird bei Einhaltung der in den bauaufsichtlichen 

Zulassungen festgelegten Verankerungstiefe eingehalten. 

RESUME 

Les vis d'ancrage sont un système de ancrage relativement nouveau. Leur 

principal avantage est une installation rapide et facile. Un trou est foré dans le 

béton et les spires sont taraudées dans le béton par la vis lors de sa mise en 

place. Les vis d'ancrage transfèrent les charges de tension dans le béton par le 

couplage mécanique des spires. En raison de leur mécanisme porteur, les vis 

d'ancrage avec un agrément technique du DIBt peuvent être utilisées pour des 

ancrages dans le béton fissuré et non-fissuré. Le mécanisme de rupture pour les 

vis d'ancrage est la rupture par cône de béton. Une augmentation de la 

profondeur d'encrage est accompagnée d'une diminution du rapport de la 

profondeur du cône de béton à la profondeur d'encrage. La charge de rupture des 

vis d'ancrage à filetage continu sur toute la profondeur d'ancrage augmente 

proportionnellement à hef 1,5 (hef = profondeur d'ancrage effective), elle est 

28


néanmoins environ 20 % inférieure à la charge de rupture des chevilles à 

expansion et des chevilles à verrouillage de forme avec la même profondeur 

d'ancrage. Afin que les vis d'ancrage puissent fonctionner correctement, les 

filetages taraudés dans le béton ne doivent pas être endommagés pendant 

l'installation. Ceci est réalisé si l'on respecte la profondeur d'ancrage définie 

dans l'agrément technique. 

KEYWORDS: concrete screw, shearing-off of threads, mechanical interlock 






In Germany there are currently three different types of concrete screws 

from three manufacturers approved by the DIBt for fastenings with single 

anchors and groups in cracked and non-cracked concrete [1,2,3]. Further 

technical approvals exist for suspended ceilings and other comparable static 

systems. 

During the technical approval process a large number of tests were 

conducted at the Institute of Construction Materials at the University of 

Stuttgart. Furthermore, the load bearing behaviour of concrete screws was 

systematically investigated through experimental and numerical studies within 

the scope of a research project. Important results of research reports [5, 6, 7, 8] 

are presented below. 

2. CONCRETE SCREWS WITH TECHNICAL APPROVAL OF THE 

DIBT 

Figure 1 shows three concrete screws with a technical approval by the 

DIBt. The screws are intended for a drill hole diameter of d0 = 10mm and are 

made of galvanised steel. They differ principally in steel strength, core diameter 

and thread geometry. Two of the concrete screws have small steel teeth at the 

end of the screw for cutting the threads into the concrete. The third concrete 

screw has alternating high and low screw threads. Grooves are cut into the 

concrete by the specially formed high screw threads. 

29 



Figure 1: Concrete screws (d0 = 10mm) with a technical approval of the DIBt 

Concrete screws made of galvanised steel intended for a drill hole diameter 

of d0 = 5mm and d0 = 6mm are approved for suspended ceilings. Concrete 

screws with a drill bit diameter of d0 = 8mm and d0 =10mm have a technical 

approval for the fastenings of statically determined and undetermined supported 

components in cracked and non-cracked concrete. Fastenings with single 

anchors and groups are allowed. 

Technical approvals also exist for concrete screws made of stainless steel 

with drill bit diameters of d0 = 6mm to d0 = 10mm. To aid in the cutting of 

threads into the concrete, one concrete screw has an end made of galvanised 

steel. This end cannot be added to the embedment depth. Another concrete 

screw has small cutting pins made of carbon steel in the first turns to cut the 

threads into the concrete. 

While concrete screws made of galvanised steel are only allowed for use in 

dry environments, the concrete screws made of stainless steel can be used 

outdoors, in industrial environments and near the sea. 

Concrete screws made of galvanised steel are cold-rolled and subsequently 

tempered and heat-treated. Residual stress and incipient cracks in the steel can 

result from this process. To insure flawless products, special tests must be 

carried out during manufacturing within the scope of the internal quality control. 

Concrete screws made of galvanised steel, which are produced according to 

requirements for the technical approvals, have an indefinite lifespan in dry 

environments. If concrete screws made of galvanised steel are used in 

environments with a high corrosion risk (e.g. outdoors), a brittle failure can 

occur as a consequence of stress corrosion cracking. The time until failure 

cannot be predicted. In these cases concrete screws made of stainless steel (or 

other types of fastenings) must be used. 

30


In the following section results of tests with the concrete screws type 1 to 

type 3 are presented. It is pointed out that the numbering of the concrete screw 

types is not the same as shown in Figure 1 or in the cited references. 

3. LOAD BEARING BEHAVIOUR OF CONCRETE SCREWS 

During installation, concrete screws cut a thread into the wall of the drilled 

hole (Figure 2). Therefore, tensile loads are transferred into the base material by 

diagonal struts, i.e. mechanical interlock (Figure 3a). The load transfer 

mechanism is similar to that of deformed reinforcing bars cast into concrete 

(Figure 3b) because the flanks of the screw thread function in a similar manner 

as the ribs of reinforcing bars. However, the laws for deformed reinforcing bars 

are only partially valid for concrete screws. One reason for this is that damage 

due to small outbreaks in the threads cut into the wall of the drilled hole can 

occur, which reduce the area for the mechanical interlock. Additionally, the core 

diameter of the concrete screw is smaller than the drill hole diameter to allow for 

easier installation. Consequently, the lateral restraint of the concrete is lost in the 

region of the highly loaded concrete consoles. To achieve sufficient load transfer 

into the concrete, the „relative rib area“ of concrete screws, which corresponds 

roughly to the ratio between the depth and the spacing of the threads cut into the 

wall of the drilled hole, is much larger than that of commercially available 

deformed reinforcing bars. 

Figure 2: Concrete screw and a thread cut into the wall of the drilled hole [9] 

31 



a) 

b) 

Figure 3: Transmission of tension load into concrete 

a) Concrete screw 

b) Cast-in-place deformed reinforcing bar 

4. INSTALLATION OF CONCRETE SCREWS 

Concrete screws are normally screwed into the concrete using an electricscrew-gun. 

In technical approvals the power class [2,3] or the type of electricscrew-gun 

[1] is specified. The threads cut into the concrete must not be 

destroyed during installation. Limiting the applied torque can do this. Concrete 

screws can also be screwed in with a torque wrench. It cannot be excluded that 

concrete screws should not be screwed in using a commercial screw-wrench, 

because the torque necessary for tightening up after the screw head reaches the 

attachment can range between wide limits and therefore the threads cut into the 

concrete might be destroyed. 

The necessary installation torque for cutting the threads into the concrete 

should be small in order to achieve an easy installation. Moreover, the resistance 

against shearing-off of the threads should be as high as possible, so that the 

threads cut into the concrete are not destroyed while tightening up the concrete 

screws. 

Figure 4 shows the measured torques while screwing in a concrete screw 

(drill bit diameter d0 = 8mm) dependent on the swing angle. The failure 

happened by shearing-off of the threads. The anchorage material consisted of 

fine-grained concrete (maximum aggregate size 8mm) of the strength class B25. 

32

Drehmoment [Nm] 

125 

100 

75 

50 

25 


0 

0 250 500 750 1000 1250 1500 1750 2000 

Drehwinkel [Grad] 

Figure 4: Typical relationship between torque moment and swing angle (Concrete B25, 

grading curve BC 8, d0= 8mm, Failure mode: Shearing-off of the thread [10] 

Before the screw head reached the attachment, the necessary installation 

torque varied only slightly. If the concrete contains coarser aggregates, torque 

peaks can occur if a thread is cut into a big piece of aggregate. 

After the screw head reaches the attachment, the torque on the concrete 

screw rises sharply to the peak value TD. Subsequently, the shearing-off of the 

threads begins and the torque decreases rapidly to zero. The damage to the 

concrete threads after overtightening the concrete screw is shown in Figure 5. 

Figure 5a shows the threads after the screw head reaches the attachment 

(installation torque TE). For comparison, the threads cut into the concrete by the 

concrete screw (d0 = 10mm) at the remaining torques of TRest ~ 0,75TD,m and 

TRest ~ 0,19TD,m after reaching the peak value TD,m are shown in Figure 5b and 

Figure 5c, respectively. 

a) b) c) 

Figure 5: Threads cut into the wall of the drilled hole, concrete screw type 2 [10] 

a) Tinst = TE 

b) TRest = 100Nm (~0,75 TD,m) 

c) TRest = 25Nm (~0,19 TD,m) 

33 



again with an electric-screw-gun, the minimum time until failure tK decreases in 

comparison with concrete screws that were not unscrewed. If the unscrewing of 

the concrete screw takes place with an electric-screw-gun, the time until failure 

tK decreases significantly because it is not possible to unscrew concrete screws 

in a controlled manner with an electric-screw-gun. 

t K [sec] 

16 

12 

8 

4 

0 

installation with 

electric-screw-gun 

installation /unscrewing / 

installation with electricscrew-gun 

installation with electric-screwgun, 

unscrewing with torque 

wrench, installation with electricscrew-gun 

Figure 12: Influence of unscrewing of concrete screws on the time tK until failure (d0 = 10mm, 

fcc = 26N/mm², grading curve BC8, dcut = 10,44mm, hnom = 60mm, electric-screw-gun 1) 

4.3 Remaining load-carrying capacity 

To investigate the influence of the installation torque, i. e. the overtightening 

of the concrete screw, on the pull-out failure load, the concrete screws 

were installed until the screw head reached the attachment (T = TE), prestressed 

with T ~ 0,9 TD,m or until the torque moment fell to a preset value T = TRest after 

reaching the maximum torque. Afterwards the concrete screws were pulled out. 

Figure 13 shows the measured failure loads depending on the installation torque. 

If the torque of the concrete screw is stopped immediately after reaching the 

maximum torque, the measured pull-out failure loads are in the same range like 

in the tests with concrete screws that were prestressed with T = TE or with 

T ~ 0,9 TD,m. Furthermore, the load-displacement behaviour does not differ 

significantly (Figure 14). If the concrete screws are turned further, the failure 

load falls rapidly, because the threads cutting into the wall of the drilled hole are 

destroyed (cp. Figure 5). Furthermore, the load-displacement behaviour is less 

favourable. The behaviour shown in Figure 13 and Figure 14 also applies to 

other types of concrete screws if they are seated with an embedment depth at 

which shearing-off of the threads is possible. 

38

Nu [kN] 

16 

14 

12 

10 

8 

6 

4 

2 

0 

1,4 T = TE before shearing 

0,9 1,2 

TU,m = 135Nm 

1 


0,8 

T/T U,m [-] 

after reaching TU,m 

Figure 13: Influence of torque before and after reaching the failure torque on the pull-out 

load (d0 = 10mm, hnom = 50mm, dcut = 10,42 mm, fcc = 30N/mm²) 

Nu [kN] 

16 

14 

12 

10 

8 

6 

4 

2 

T = 0,19 TD,m 

T = 0,75 TD,m 

0 

0 2 4 6 8 10 

s [mm] 

0,6 

0,4 

0,2 

T = TE 

T = 0,75xTD,m 

T = 0.19xTD,m 

Figure 14: Influence of the torque moment before and after reaching the 

failure torque on the load-displacement curves 

4.4 Required embedment depth 

In practice it cannot be excluded that concrete screws are further tightened 

after the screw head reaches the attachment, e. g. if the electric-screw-gun is not 

stopped immediately or if the attachment should be tightened against the surface 

of the concrete slab with a standard screw-wrench. Unscrewing of the concrete 

screws and screwing them in again can also occur. This may damage the threads 

cut into the wall of the drilled hole, if the embedment depth is not deep enough 

because in practice it is normally not possible to stop the installation after 

reaching the maximum torque TD. This has been shown by experiences in 

practice. A check of concrete screws (d0 = 6mm) that were seated with a small 

embedment depth showed that shearing-off of the threads during the installation 

had occurred with about 15% of the screws. 

39 

0 



Table 1: Characteristic values for the resistances under tension load of concrete screws 

(d0 = 10mm) with a Technical Approval of the DIBt 

Type of concrete screw [1] [2] [3] 

Drill hole diameter d0 [mm] 10 10 10 

Embedment depth hnom [mm] 70 75 85 

Steel failure 

Characteristic resistance NRk,s [kN] 54,1 75,4 58 

Characteristic resistance in noncracked 

concrete B 25 

Characteristic resistance in 

cracked concrete B 2 

Nominal effective embedment 

depth 

Pull-out failure 

NRk,p [kN] 12,0 16,0 20,0 

NRk,p [kN] 7,5 12,0 12,0 

Concrete cone failure 

hef,cal [mm] 50 50 60 

Characteristic screw spacing scr,N [mm] 150 150 180 

Characteristic edge distance ccr,N [mm] 75 75 90 

8. SUMMARY 





Concrete screws transfer tensile loads into the base material by mechanical 

interlock of the threads. Due to their load-bearing mechanism, concrete screws 

with a technical approval of the DIBt can be used for fastenings in cracked and 

non-cracked concrete. 

The typical failure mechanism for concrete screws is concrete-cone failure. 

With increasing embedment depth the ratio of the depth of the concrete failure 

cone to the embedment depth decreases. The failure load of concrete screws 

with continuous threads along the entire embedment depth increases 

proportionally to hef 1,5 (hef = effective embedment depth), but it is about 20 % 

49 



smaller than the failure load of expansion and undercut anchors with the same 

embedment depth. 

In order for concrete screws to function properly, the threads cut into the 

wall of the drilled hole must not be damaged during the installation. This 

requirement is achieved by using the embedment depth defined in the Technical 

Approvals. 

9. ACKNOWLEDGMENT 

The primary funding for this research was provided by the Adolf Würth 

GmbH & Co. KG. The support of this manufacturer is very much appreciated. 

Special thanks are also accorded to Beate Vladika and Matthew Hoehler who 

spent many hours in improving the English. 

10. REFERENCES 

[1] [Deutsches Institut für Bautechnik] Allgemeine Bauaufsichtliche 

Zulassung Z-21.1-1712 für Hilti Schraubanker HUS-H zur Verankerung 

im gerissenen und ungerissenen Beton, Berlin, 2001 


Zulassung Z-21.1-1549 für HECO-MULTI-MONTI-Schraubanker MMS 

zur Verankerung im gerissenen und ungerissenen Beton, Berlin, 2001 


Zulassung Z-21.1-1624 für Toge Betonschraube TSM zur Verankerung 

im gerissenen und ungerissenen Beton, Berlin, 2001 

[4] European Organisation for Technical Approvals (EOTA): Leitlinie für die 

europäisch-technische Zulassung von Metalldübeln zur Verankerung in 

Beton. Deutsches Institut für Bautechnik, 28. Jahrgang, Sonderheft Nr. 16, 

Berlin, Dezember 1997 

[5] Küenzlen, J. H. R.; Eligehausen, R.: Setz- und Ausziehversuche in 

ungerissenem Beton mit Schraubdübeln. Bericht Nr. AF01/01-E00202/1, 

Institut für Werkstoffe im Bauwesen, Universität Stuttgart, 2001, nicht 

veröffentlicht 

[6] Küenzlen, J. H. R.; Eligehausen, R.: Tragverhalten von Schraubdübeln in 

niederfestem Beton. Bericht Nr. W8/1-01/1, Institut für Werkstoffe im 

Bauwesen, Universität Stuttgart, 2001, nicht veröffentlicht 

50


[7] Küenzlen, J. H. R.; Eligehausen, R.: Tragverhalten von Schraubdübeln in 

niederfestem Beton. Bericht Nr. W8/3-01/3, Institut für Werkstoffe im 

Bauwesen, Universität Stuttgart, 2001, nicht veröffentlicht 

[8] Küenzlen, J. H. R.; Eligehausen, R.: Einfluss verschiedener Parameter auf 

die Höchstlasten von Schraubdübeln, Institut für Werkstoffe im 

Bauwesen, Universität Stuttgart, 2001, Bericht in Vorbereitung 

[9] Küenzlen, J. H. R.; Sippel, T. M.: Behaviour and Design of Fastenings 

with Concrete Screws. In: RILEM Proceedings PRO 21 „Symposium on 

Connections between Steel and Concrete“, Cachan Cedex, 2001, S. 919- 

929. 

[10] Küenzlen, J. H. R.: Drehmomentversuche mit Schraubdübeln in 

ungerissenem Beton. Jahresbericht 2000/2001, Institut für Werkstoffe im 

Bauwesen, Universität Stuttgart, 2001 

[11] Eligehausen, R.; Hofacker, I. N.; Spieth, H. A.; Küenzlen, J. H. R.: Neue 

Entwicklungen in der Befestigungstechnik, Tagungsband, IBK-Bau- 

Fachtagung 263: Dübel und Befestigungstechnik, 2000, S. 2.1-2.14, 

[12] Meszaros, J.,: Tragverhalten von Einzelverbunddübeln unter zentrischer 

Kurzzeitbelastung. Dissertation, Universität Stuttgart, 2001 

[13] Eligehausen, R.; Fuchs, W.; Mayer, B.: Tragverhalten von 

Dübelbefestigungen bei Zugbeanspruchung. Beton + Fertigteil-Technik 

1987, Heft 12, S. 826-832 und 1988 Heft 1, S. 29-35. 

[14] Eligehausen, R.; Mallée, R.: Befestigungstechnik im Beton- und 

Mauerwerkbau. Ernst & und Sohn, Berlin, 2000. 

[15] Fuchs, W.; Eligehausen, R.: Das CC-Verfahren zur Berechnung der 

Betonausbruchlast von Verankerungen. Beton- und Stahlbetonbau, 1995, 

Heft 1, S. 6-9, Heft 2, S. 38-44, Heft 3, S. 73-76. 

[16] Deutsches Institut für Bautechnik: Bemessungsverfahren für Dübel zur 

Verankerung in Beton (Anhang zum Zulassungsbescheid). Berlin, 1993 

[17] DIN 1045, Beton und Stahlbeton, Bemessung und Ausführung, Ausgabe 

1978 

51 



52

Prestressed hollow-core concrete slabs – Problems and possibilities in fastening techniques 

PRESTRESSED HOLLOW-CORE CONCRETE SLABS – PROBLEMS 

AND POSSIBILITIES IN FASTENING TECHNIQUES 

SPANNBETON-HOHLDECKENPLATTEN – PROBLEME UND MÖG- 

LICHKEITEN IN DER BEFESTIGUNGSTECHNIK 

DALLES ALVEOLAIRES EN BÉTON PRÉCONTRAINT – PROBLÈ- 

MES ET POSSIBILITÉS DES TECHNIQUES D'ANCRAGE 

Clemens Lutz 

SUMMARY 

In the present, prestressed hollow-cored concrete slabs are tendentiously 

used as ceiling systems. Therefore, fastening techniques with regard to these 

slabs gain in an increasing importance. In this article, advantages of these members 

and problems by application of anchors are described, and different structural 

responses between several types of anchors are experimentally determined. 

Accordingly, it seems to be essential to choose or to adapt suitable anchors for 

ceiling systems. 


Spannbeton-Hohlplatten finden als Deckensystem eine immer breitere 

Verwendung und einen immer größeren Anwendungsbereich. Somit gewinnt 

auch eine korrekte Befestigung in diesen Platten zunehmend an Bedeutung. In 

diesem Artikel wird auf die Vorteile der Platten, aber auch auf die Problematik, 

die bei der Montage von Dübeln entstehen, eingegangen. Ferner wird gezeigt, 

dass es große qualitative Unterschiede bezüglich der Tauglichkeit verschiedener 

Befestigungssysteme gibt, weshalb eine sorgfältige Auswahl bzw. Anpassung 

prinzipiell geeigneter Dübel stattfinden muss. 

RESUME 

Actuellement, les dalles alvéolaires en béton précontraint sont utilisées de 

plus en plus fréquemment. Par conséquent, les ancrages appropriés gagnent 

d'importance. Dans cet article, nous traitons les avantages de ces dalles et les 

problèmes reliés à l'utilisation de chevilles. De plus, nous montrons que les dif- 

53 


C. LUTZ 

férents systèmes révèlent de grandes différences qualitatives, et qu'il est par 

conséquent essentiel de choisir et d'adapter des ancrages adéquats. 

KEYWORDS: prestressed hollow-core concrete slab, ceiling, anchorageable thickness, 

anchor 

1. ADVANTAGES AND PROBLEMS 

Prestressed hollow-cored concrete slabs made of high-strength concrete are 

prefabricated concrete members with large hollow proportions. In practice, they 

are interconnected after assembly by joint grouting compound. In comparison 

with conventional concrete members, this type of concrete plates has a lot of 

economical advantages, especially in saving material, energy and in reducing 

weight of transportation. Outstanding features are quality control, schedule time 

and costs. Additionally, formworks which are used to produce in-situ concrete 

are saved in application of these slabs. In the present, this ceiling system is increasingly 

used in industrial buildings, office buildings and also in domestic architecture. 

Figure 1 shows cross sections of two types of prestressed hollowcored 

concrete slabs (with different minimal anchorageable material thickness: 

25 mm and 30 mm). 

Figure 1: Cross sections of two types of prestressed hollow-cored concrete slabs 

(1: cavity, 2: prestressed wire, 3: steel, 4: minimal anchorageable material 

thickness dmat, here: 25 mm and 30 mm) 

54 

1 

4 

4


In spite of above mentioned advantages, the application of anchors in 

prestressed hollow-cored concrete slabs is not satisfied, particularly in case of 

thin members. The worst case is that anchors are fastened in near of position A 

(see fig. 2). The distance between the opposite of casting side and the hollows is 

the smallest. This distance is defined as minimal anchorageable material thickness 

dmat (in German: Spiegeldicke). For some types of slabs the value dmat is 

very small. This small value of thickness is relevant for load carrying capacity of 

anchor systems. Tables 1 and 2 show experimentally measured thickness dmat of 

slabs with a minimal anchorageable thickness of 30 mm and 25 mm respectively. 

All values in table 1 are above 30 mm. Some values in table 2 are only 

just 25 mm. For slabs with dmat = 25 mm there is no sufficient reserve in comparison 

to slabs with dmat= 30 mm! Furthermore, crashing of concrete closed to 

the hollows often occurs during drilling. Consequently, the minimal anchorageable 

material thickness and also the effective anchorage depth for anchors are 

reduced (see fig. 3) and load carrying capacities of ceiling systems are negatively 

influenced. Therefore, it is necessary to determine whether all types of 

fasteners are suitable to be used in prestressed hollow-cored concrete slabs. Additionally, 

it is prohibited to install an anchor in near of a strand of wire because 

of interests of safety (zone C in fig 2). 

zone C zone C 

zone B 

zone A 

zone B 

Figure 2: Sectors of a prestressed hollow-cored concrete slab. 

Zone A: minimal anchorageable material thickness; 

Zone B: anchorageable sector; 

Zone C: prohibited sector for fastenings because of interests of security 

(prestressed concrete wire) 

55 


C. LUTZ 

Table 1: measured values dmat (slab with a minimal anchorageable thickness of 30 mm) 

dmat [mm] (measured minimal 

anchorageable material 

thickness) 

39,6 42,3 

43,1 44,2 

42,3 45,0 

40,9 43,8 

37,8 39,2 

Range [mm] 38 (>30) to 45 

Average [mm] 41,8 

Variation coeff. [%] 5,70 

Table 2: measured values dmat (slab with a minimal anchorageable thickness of 25 mm) 

dmat [mm] (measured minimal 

anchorageable material 

thickness) 

25,1 26,7 

26,7 28,5 

27,6 28,5 

26,8 27,5 

26,2 28,6 

25,0 26,1 

Range [mm] 25 to 29 

Average [mm] 26,9 

Variation coeff. [%] 4,60 

dmat dmat, eff 

Figure 3: The minimal anchorageable material thickness dmat after drilling is reduced 

(:=dmat, eff < dmat). 

56


2. POSSIBILITIES AND TEST RESULTS 

In general, there are three types of possibilities to fasten installation pipes, 

suspendic and acoustics ceilings, lighting appliances, safety precaution systems 

and beams (see fig. 4). In the following, fastenings with different types of anchors 

according to possibility (1) will be studied in detail. As results, these anchors 

applied in prestressed hollow-cored concrete slabs show their quite different 

suitability. 

There are already several types of special fasteners on the market, which 

have approvals for using in hollow-cored concrete slabs. Objective of this work 

is to investigate suitability and quality of other types of anchors in these members. 

Therefore, pull-out tests were carried out in the uncracked concrete zone of 

these slabs with a minimal anchorageable material thickness dmat of 25 mm and 

30 mm. Herein, different types of anchors – concrete screws, injection anchors, 

suspendic ceiling fasteners, deformation-controlled expansion anchors and 

torque-controlled expansion anchors – were used. The sizes of anchors chosen 

for these experiments were between M6 and M10. 

(1) only anchors 

(2) post-installed bonded rebar connections; concrete suspension 

(3) construction, fastening through the slab 

Figure 4: Three types of possibilities to fasten installation pipes, suspendic and 

acoustics ceilings, lighting appliances, safety precaution and beams 

57 


C. LUTZ 

For each test, one borehole was produced with the help of a hammer drill. 

Position of the borehole was chosen in such a way that the thickness of concrete 

corresponds to the minimal anchorageable thickness dmat. Depth of the borehole 

is equal to this minimal thickness (position A, fig. 2). Typical crashing of concrete 

closed to the hollows was often observed after drilling. Consequently, the 

effective anchorage depth was reduced. After installation of the anchor system 

the fastener was subjected to concentric tension up to failure. For concrete 

screws, setting tests with concrete screws were also carried out additionally [1]. 

Figure 5 outlines an equipment for pull-out tests where one load cell, two 

LVDTs and a steel support frame are used. Figure 6 shows a pull-out cone of a 

concrete screw. Pull-out test results for different types of anchors are represented 

in following figures. Figure 7 shows measured load carrying capacities of 

different types of anchors used in concrete slabs with a minimal anchorageable 

material thickness of 30 mm. All test results are given in relation to the failure 

load of concrete screws, type 1 (which is chosen as reference anchor). 

Figure 5: Pull-out tests in a prestressed hollow-cored concrete slab 

58


Figure 6: Pull-out cone of a concrete screw 

From figure 7, it can be seen that the highest load carrying capacity was obtained 

in torque-controlled expansion anchors (column 7). The average failure 

load was 93% higher in comparison to anchor, type 1 (reference anchor). Relative 

loading carrying capacities of other anchors are summarized in the following: 

- concrete screws, type 2 (column 2) +64% 

- deformation-controlled expansion anchors (column 4) +30% 

- injection anchors (column 6) +12% 

- special fasteners for hollow-cored concrete slabs (column 5) ±0 

- concrete screws, type 1 (column 1) ±0 

- suspendic ceiling fasteners (column 3) −26% 

59 


C. LUTZ 

Figure 7: 

Figure 8: 

Relative values of averaged 

failure load Nu.m [%] 

Nu,m [%] 

250 

200 

150 

100 

50 

0 

100 

164 

74 

130 

100 

112 

193 

1 2 3 4 5 6 7 

Types Dübelart of anchors 

Pull-out test results for different types of anchors (minimal anchorageable material 

thickness: 30 mm). All test results are given in relation to the failure load of 

a concrete screw, type 1 [1]. 

1: concrete screws, type 1 


3: suspendic ceiling fasteners 

4: deformation-controlled expansion 

anchors 

Relative values of averaged 

failure load Nu.m [%] 

Nu,m [%] 

350 

300 

250 

200 

150 

100 

50 

0 

100 

193 

5: special fasteners for hollow-cored 

concrete slabs 

6: injection anchors 

7: torque-controlled expansion anchors 

120 

289 

1 2 3 4 

Types Dübelart of anchors 

Pull-out test results for different types of anchors (minimal anchorageable material 

thickness: 25 mm). All test results are given in relation to the failure load of 

a concrete screw, type 1. 


2: deformation-controlled expansion 

anchors 

60 

3: suspendic ceiling fasteners 

4: special fasteners for hollow-cored 

concrete slabs


Figure 8 shows experimental results of different types of anchors used in a 

thin hollow-cored concrete slab with a minimal anchorageable material thickness 

of 25 mm. All mean values of load carrying capacities are represented in 

relation to the averaged failure load of anchor, type 1. Relative loading carrying 

capacities of other types of anchors are summarized as follows: 

- special fasteners for hollow-cored concrete slabs (column 4) 

+189% 

- deformation-controlled expansion anchors (column 2) +93% 

- suspendic ceiling fasteners (column 3) +20% 

- concrete screws, type 1 (column 1) ±0 

For most of structural designs the averaged load carrying capacity is not 

alone the value which characterizes the material properties. Displacements and 

statistic values are also important factors. In figure 9 load-displacementdiagrams 

on the left and right are compared (diagram a and b): Failure load and 

statistic values according to the load carrying capacities of these both types of 

anchors are almost the same (see also table 3 for statistic values), whereas the 

displacements and statistic values according to the displacements are quite different. 

Therefore, it may be questioned, which type of anchor is more suitable 

for hollow-cored concrete slabs. Anchors of type 1 behave more brittle, anchors 

of type 2 behave more ductile. In this case, displacement at the permissible load 

is essential. Type 1 seems to behave more positive than type 2. In figure 9 c.) 

and d.) anchors of type 3 reach higher load carrying capacities on average in 

comparison with type 4, but they have also higher displacements. It is harmful if 

displacements are to high and come outside of linear area (see fig. 10). 

61 


C. LUTZ 

load 

load 

0 

Nx 

0 

Ny 

displacement 

a.) type 1 

displacement 

load 

0 0 

sx 

load 

displacement 

b.) type 2 

displacement 

c.) type 3 d.) type 4 

Nx 

0 

Ny 

0 0 

0 

sy sy 

Figure 9: Load-displacement-curves for different types of anchors. 

Diagrams a.) and b.): Difference in displacement with almost the same failure load. 

Diagrams c.) and d.): Difference in failure load and displacement 

Table 3: Statistic values of two types of anchors in a prestressed hollow-cored concrete slab 

with dmat= 30 mm (see figure 9) 

a.) Type 1 b.) Type 2 

Variation coefficient at failure load Nu,m [%] 17 17 

Displacement at 0,5 Nu,m [%] 100 560 

Variation coefficient for displacement at 0,5 Nu,m [%] 11 25 

62 

sx


Obviously, the load carrying capacity is not only the aspect characterizing 

anchors. In principle, three groups of anchors could be distinguished according 

to their load-displacement-behaviours (see fig. 11): 

- Group 1: anchors with little displacements, i.e. the tested injection anchors. 

- Group 2: anchors with larger displacements and high load carrying capacities, 

i. e. one of the tested torque-controlled expansion anchors. 

- Group 3: anchors with larger displacements and reduced load carrying capacities. 

Load N 

Linear-elastic 

area 

Linear-elastic area 

0 0.5 1 1.5 2 

Displacement s [mm] 

Figure 10: Experimental load-displacement-curves and simplified linear curves for two types 

of anchors in a prestressed hollow-cored concrete slab (here: with dmat= 30 mm). 

63 


C. LUTZ 

Load N [kN] 

Group 1 

Displacement s [mm] 

Group 2 

Group 3 

Figure 11: Load-displacement-curves of anchors in prestressed hollow-cored concrete slabs 

(here: with dmat= 30 mm) can be distinguished in three groups. There are serious differences 

in carrying load capacities, but also in displacements 

According to the above mentioned observations, it can be concluded that 

anchors of group 1 in connection with serious statistic values in according to 

load carrying capacities and displacements are most suitable to apply in 

prestressed hollow-cored concrete slabs. Though, further tests have to be done 

(sustained load, fatigue tests and so on). 

REFERENCES 

[1] LUTZ, C.: Anchors in prestressed hollow-cored concrete slabs. IWB 

Activities 3 (2001) 

[2] LUTZ, C.: Nachträgliche Befestigungen in Spannbeton-Hohlplatten. IWB 

Mitteilungen, Jahresbericht 2000-2001 

64

Pore-size determination from penetration tests on concrete with n-decane 

PORE-SIZE DETERMINATION FROM PENETRATION TESTS ON 

CONCRETE WITH N-DECANE 

PORENGRÖSSENBESTIMMUNG AUS N-DECAN-EINDRING- 

VERSUCHEN IN BETON 

DETERMINATION DES PORES DU A LA PENETRATION DE 

N-DECANE EN BETON 

Hans W. Reinhardt, Arno Pfingstner 

SUMMARY 

Absorption and infiltration tests on concrete mixes have been carried out 

with n-decane. The test results show a good agreement with theoretical predictions. 

The results indicate that the main parameter on the penetration is the water-cement 

ratio. Pore sizes which are reached are different for the absorption 

test and the infiltration test. 


Absorptions- und Infiltrationsversuche wurden an verschiedenen Betonen 

mit n-Decan durchgeführt. Die Ergebnisse haben eine gute Übereinstimmung 

mit theoretischen Vorhersagen gezeigt. Die Ergebnisse lassen den Schluss zu, 

dass der Wasserzementwert die maßgebliche Größe für das Eindringverhalten 

ist. Porengrößen, die erreicht wurden, sind bei Absorptions- und Infiltrationsversuchen 

verschieden. 

RESUME 

Des essais d'absorption et d'infiltration ont été réalisés sur des bétons de 

différentes compositions avec du n-décane. Les résultats montrent une bonne 

concordance avec des prévisions théoriques. Les résultats indiquent que le paramètre 

principal pour la pénétration dans le béton est le rapport eau-ciment. Les 

tailles des pores atteintes sont différentes pour les essais d'absorption et les essais 

d'infiltration. 

KEYWORDS: Concrete, n-decane, penetration, absorption, infiltration, highperformance 

concrete 

65 


H.W. REINHARDT, A. PFINGSTNER 

MOTIVE 

Since several years, tests have been carried out on the penetration behaviour 

of organic fluids in concrete [1-5]. Two types of tests are being carried out: 

the capillary suction test and the infiltration test with a certain hydraulic head. 

The test results show that the capillary suction test satisfies usually technical 

requirements for the application of the material properties in assessing the behaviour 

of real structures. However, from a scientific point of view both tests 

reveal more than the material property only. Comparing the test results of both 

test methods one can calculate the pore radius of capillary pores in concrete. 

This will be shown in the following. 

EXPERIMENTAL SET UP 

For suction tests, specimens were placed into the test fluid. The samples 

rested on glass rods to allow free access of the testing fluid to the inflow surface. 

The fluid level was approx. 10 mm above the lower end of the specimen. Penetration 

occurred by capillary forces acting against gravity. 

The experimental set-up for infiltration tests described in DAfStb guideline 

[6] was slightly modified. Preliminary tests had proven that the pressure 

head of 40 (+/- 5) cm specified there was too small to obtain measurable differences 

to capillary suction tests for high performance concretes. The required external 

pressure was estimated by calculation from the pore radius distributions of 

comparable concretes and fixed to 0,2 bar (20 kN/m 2 ) for all infiltration tests. 

This pressure was produced by a nitrogen bottle connected to the funnels on the 

samples by tubes. 

66

CONCRETES USED 


The composition of the concretes used is given in Table 1. Table 2 shows 

some relevant properties. The air content has been measured in the fresh state. 

Table 1. Composition of concretes 

Concr. Aggregates 

[kg/m 3 Grading Cement 

] 

[kg/m 3 Type of cement 

Wadded 

] 

[kg/m 3 SF (solid) 

] [kg/m 3 RE 

] [kg/m 3 SP 

] [kg/m 3 Wtot./(C+SF) 

] - 

MR 1905 AB 16 320 CEM I 32,5 R 160 0 0 2.50 0.50 

M1 1822 AB 16 338 CEM I 32,5 R 186 0 0 0.80 0.55 

M2 1535 AB 2 467 CEM I 32,5 R 257 0 0 0 0.55 

M3 1882 AB 32 309 CEM I 32,5 R 170 0 0 0 0.55 

M4 1895 U 16 309 CEM I 32,5 R 170 0 0 0 0.55 

M5 1677 C 16 405 CEM I 32,5 R 223 0 0 0.50 0.55 

M6 1769 AB 16 485 CEM I 42,5 R 150 0 0 9.30 0.32 

M7 1762 AB 16 465 CEM I 42,5 R 126 20 0 7.00 0.31 

M8 1755 AB 16 445 CEM I 42,5 R 109 40 2.56 10.80 0.32 

M9 1748 AB 16 425 CEM I 42,5 R 89 60 2.56 12.00 0.32 

M10 1441 AB 2 615 CEM I 42,5 R 153 55 3.59 11.08 0.32 

M11 1813 AB 16 338 CEM III/B 186 0 0 0 0.55 

M11 1522 AB 2 467 CEM III/B 257 0 0 0 0.55 

SF: silica fume RE: retarder SP: plasticizer C: cement W: water 

Table 2. Some properties of the concretes tested, mean of three tests 

Properties of fresh concrete Compressive strength after 28 days 

workability 1) , density of air con- density of hard. 

flow fresh concrete tent concrete 2) 

Mix 

Compressive strength 

smallest value mean value 

[cm] [kg/dm 3 ] [%] [kg/dm 3 ] [N/mm 2 ] [N/mm 2 ] 

MR/1 41.8 2.34 1.8 2.35 52.3 53.8 

MR/2 44.8 2.33 1.2 2.36 51.5 53.2 

M1/1 46.5 - - 2.33 41.8 44.0 

M1/2 46.5 2.35 1.5 2.33 45.4 46.2 

M2 43.5 2.16 3.6 2.19 41.8 43.1 

M3/1 44.5 2.39 0.9 2.37 42.5 43.6 

M3/2 46.5 2.39 0.4 2.35 42.2 43.0 

M4 46.5 2.40 0.3 2.38 47.5 49.0 

M5 43.8 2.28 1.8 2.18 38.2 38.8 

M6 43.0 2.40 1.75 2.39 75.2 77.2 

M7 48.3 2.37 1.4 2.41 80.9 84.3 

M8/1 42.0 2.37 1.5 2.40 88.0 90.5 

M8/2 44.8 2.37 0.7 2.40 85.1 86.2 

M9 44.0 2.38 1.6 2.38 85.4 88.9 

M10 44.5 2.22 2.8 2.24 77.2 78.9 

M11/1 47.5 2.36 0.55 2.35 40.9 43.0 

M11/2 46.5 2.35 0.44 2.34 45.5 46.0 

M12 51.0 2.36 1.4 2.19 33.8 35.6 

1) 

workability: average diameter of the spread concrete determined by the German flow table test 

2) determined on 100 mm cubes 

69 



The density is the dry density after 28 days. The compressive strength has 

been measured on 100 mm cubes (150 mm for M3, due to the maximum aggregate 

size of 32 mm) after 1 day kept in the mould, 6 days in the moist room and 

21 days in the constant climate room at 20°C and 65% RH. 

RESULTS 

The results show typically the absorbed amount of liquid as function of 

time up to 72 hours generated in the capillary suction test and in the infiltration 

test. Fig. 2 to 4 show on the left the results of the capillary suction test and on 

the right of the infiltration test. The test results can be presented as a straight line 

of the absorbed amount vs. square root of time. 

There are similar plots for the penetration depth which has been measured 

by visual observation at the epoxy resin covered surface of the specimens. The 

results are also shown in Fig. 5 to 7. 

Absorbed volume [l/m 2 ] 

16 

14 

12 

10 

8 

6 

4 

2 

MR1-SD 

MR3-SD 

M1/1-SD 

M1/3-SD 

M2-SD 

M3/1-SD 

M4/2-SD 

M5-SD 

MR1-SC 

M1/1-SC 

0 

0 1 2 3 4 5 6 7 8 

Square root of time [h 1 

9 

/2 

] 

Infiltrated volume [l/m 2 ] 

16 

14 

12 

10 

8 

6 

4 

2 

MR1-ID 

MR3-ID 

M1/1-ID 

M1/3-ID 

M2-ID 

M3/1-ID 

M4/2-ID 

M5-SD 

0 

0 1 2 3 4 5 6 7 8 


Fig. 2. Absorbed volume (left) and infiltrated volume (right) as function of square root 

of time: Portland cement concretes with normal strength 

70 

9 

/2 

]


7 

6 

5 

4 

3 

2 

1 

M6-SD 

M7-SD 

M8/2-SD 

M8/6-SD 

M9-SD 

M10-SD 

M8/2-SC 

0 

0 1 2 3 4 5 6 7 8 




7 

6 

5 

4 

3 

2 

1 

M6-ID 

M7-SD 

M8/2-ID 

M8/6-ID 

M9-ID 

M10-ID 

0 

9 

/2 

] 

0 1 2 3 4 5 6 7 8 



of time: Portland cement concretes with high strength 


18 

16 

14 

12 

10 

8 

6 

4 

2 

M11/1-SD 

M11/3-SD 

M12-SD 

0 

0 1 2 3 4 5 6 7 8 


9 

/2 

] 


18 

16 

14 

12 

10 

8 

6 

4 

2 

M11/1-ID 

M11/3-ID 

M12-ID 

0 

0 1 2 3 4 5 6 7 8 



of time: Blast furnace slag cement concretes 

The properties of n-decane are given in Table 3. 

Table 3. Physical values of n-decane at 20°C 

Fluid Formula Density Surface tension Dynamic viscosity 

Ratio 

( / ) 0.5 

[kg/dm 3 ] [mN/m] [mN.s/m 2 ] [m 0.5 /s 0.5 ] 

n-decane C10H22 0.73 23.9 0.88 5.21 

With those values the results have been evaluated and are presented in Table 

4. 

71 

9 

/2 

] 

9 

/2 

] 



Table 4. Pore parameters calculated from test results with n-decane 

Concrete 

Sorptivity 

l m -2 h -1/2 

penetration coefficient 

mm h -1/2 

r, from B 

µm 

cos θ = 

r, from S 

µm 

cos θ = 

So Sp Bo Bp 1 2/π 1 2/π 

MR 0.750 0.931 10.5 12.1 0.79 0.50 1.29 0.82 

M1 1.054 1.342 12.2 14.7 1.10 0.70 1.48 0.94 

M2 1.491 1.971 14.1 17.2 1.16 0.74 1.79 1.14 

M3 1.001 1.111 12.5 13.7 0.47 0.30 0.55 0.35 

M4 1.006 1.162 13.3 15.2 0.71 0.45 0.80 0.51 

M5 1.378 1.567 13.7 15.2 0.58 0.37 0.70 0.45 

M6 0.661 0.757 10.3 11.0 0.38 0.24 0.74 0.47 

M7 0.552 0.626 9.2 10.0 0.44 0.28 0.68 0.43 

M8 0.422 0.471 7.2 8.0 0.56 0.36 0.59 0.38 

M9 0.378 0.474 7.4 9.0 1.09 0.69 1.37 0.87 

M10 0.659 0.756 8.1 8.9 0.47 0.30 0.76 0.48 

M11 1.169 1.366 13.5 15.2 0.63 0.40 0.87 0.55 

M12 1.307 2.245 11.9 18.5 3.37 2.15 4.66 2.97 

Absorption depth [mm] 

130 

120 

110 

100 

90 

80 

70 

60 

50 

40 

30 

20 

10 

0 

MR1-SD 

MR3-SD 

M1/1-SD 

M1/3-SD 

M2-SD 

M3/1-SD 

M4/2-SD 

M5-SD 

MR1-SC 

0 1 2 3 4 5 6 7 8 9 

Square root of time [h 1/2 ] 

Infiltration depth [mm] 

130 

120 

110 

100 

90 

80 

70 

60 

50 

40 

30 

20 

10 

0 

MR1-ID 

MR3-ID 

M1/1-ID 

M1/3-ID 

M2-ID 

M4/2-ID 

M5-ID 

0 1 2 3 4 5 6 7 8 9 


Fig. 5. Absorption depth (left) and infiltration depth (right): Portland cement concretes 

with normal strength 


100 

90 

80 

70 

60 

50 

40 

30 

20 

10 

0 

M6-SD 

M7-ID 

M8/2-SD 

M8/6-SD 

M9-SD 

M10-SD 

0 1 2 3 4 5 6 7 8 9 



100 

90 

80 

70 

60 

50 

40 

30 

20 

10 

0 

M6-ID 

M7-ID 

M8/2-ID 

M8/6-ID 

M9-ID 

0 1 2 3 4 5 6 7 8 9 


Fig. 6. Absorption depth (left) and infiltration depth (right): Portland cement concretes 

with high strength 

72


120 

100 

80 

60 

40 

20 

M11/1-SD 

M11/3-SD 

M11/1-SC 

0 

0 1 2 3 4 5 6 7 8 



9 

/2 

] 


120 

100 

80 

60 

40 

20 

0 

M11/1-ID 

M11/3-ID 

M12-ID 

0 1 2 3 4 5 6 7 8 9 


Fig. 7. Absorption depth (left) and infiltration depth (right): Blast furnace slag cement 

concretes 

DISCUSSION 

The sorption and infiltration tests show in Fig. 2 and 7 an almost perfect 

straight line in the square root of time plot. A second general feature is that the 

infiltration results are mostly close. In the sorption results, i.e. the pressure of 20 

kPa is not important. 

Concrete mixes M1 to M5 are made with a water-cement ratio of 0.55 but 

with variations in the grading curve. Fig. 2 shows that suction proceeds the fastest 

with a maximum grain size of two millimetre and a high cement content of 

467 kg/m 3 (M2). The same is also true for the infiltration test. Also the mix with 

fine grading C16 and a cement content of 465 kg/m 3 is fast in suction but not so 

fast in infiltration. The mixes M1, M3 and M4 vary less because the cement content 

is rather similar and also grading curves are similar. The concrete mix MR 

shows the lowest suction and infiltration rates because the water-cement ratio is 

only 0.50. 

Fig. 3 contains the results of the high performance concrete with watercement 

ratios of 0.32 and typically a high cement content. Except M10 which 

has a maximum grain size of 2 mm the others have all 16 mm maximum grain 

size. There is however a variation in silica fume content. M6 and M10 have the 

fastest absorption and infiltration. the reason for that is that there is either no silica 

fume used (M6) or the cement content is very high with 615 kg/m 3 (M10). 

One should notice that the vertical scale of Fig. 3 is less than half of Fig. 2. All 

other high performance concrete mixes show smaller absorption and infiltration 

qualities. 

73 



A blast furnace slag cement has been used in the mixes of Fig. 4. The sorptive 

tests led to results which were similar to those with Portland cement and a 

water-cement ratio of 0.55 (Fig. 2). The infiltration tests on M12 which has a 

maximum grain size of 2 mm is different from the others since the infiltration 

rate is rather high. A similar result has been obtained in Fig. 2 with Portland cement. 

The absorption depth and the infiltration depth are rather similar as can be 

seen from Figs. 5 to 7. The absolute results of MR, M1 to M5 and M11 and M12 

are almost the same i. e. the influence of the grading curve is not so strong as in 

the case of the absorbed fluid volume. However, a closer look to the small variations 

reveals that the trends of grain size and cement content are the same as in 

the case of absorbed volume. 

Fig. 6 shows the smallest absorption and infiltration depth as has been expected 

since these concretes are high performance ones. 

Table 4 contains the values of the sorptivity and the penetration coefficient. 

The sorptivity is the quotient of absorbed volume per area divided by the square 

root of time. The penetration coefficient gives the penetration depth divided by 

the square root of time. Both quantities characterise physical properties of a material. 

Both material constants have been derived from sorption and infiltration 

tests, So and Sp and Bo and Bp respectively. 

The sorptivity So is in the range of 1.0 to 1.49 l m -2 h -1/2 for concrete with a 

water-cement ratio of 0.55. The corresponding value Sp lies in the range of 1.11 

to 1.97 l m -2 h -1/2 . The difference between sorption test and infiltration test is 

consistent. High performance concretes M6 to M10 show considerably lower 

values So between 0.38 and 0.66 l m -2 h -1/2 and Sp between 0.47 and 0.76 l m -2 

h -1/2 . The mixes with blast furnace slag cement fit into the ranges of mixes with 

Portland cement except M12 in the infiltration test with a high value of 2.24 l 

m -2 h -1/2 . 

The penetration coefficient Bo ranges between 12.2 and 14.1 mm h -1/2 for 

mixes with a water-cement ratio of 0.55. Bp lies between 13.7 and 15.2 mm h -1/2 , 

i. e. a slight increase due to the pressure of 20 kPa. With a lower water-cement 

ratio of 0.32 the Bo drops to 7.2 and 10.3 mm h -1/2 and Bp drops to 8.0 and 11.0 

mm h -1/2 . All results are consistent as the influence of grain size, water-cement 

ratio and pressure are concerned. The B-values for blast furnace slag cement 

concrete are similar to those of Portland cement concrete. 

74


The effective pore radius r can be calculated from Bo and Bp as shown in 

Eq. (4) or equivalently also from the sorptivities since penetration coefficient 

and sorptivity are linked together via the porosity ε (see Eq. (5)). Since the porosity 

levels out a similar equation occurs for the sorptivity as for the penetration 

coefficient. 

Table 4 contains the results. It can be seen that the pore sizes range be- 

tween about 0.2 to more than 1.0 µm when the content angle is taken to zero. 

The values decrease when the cosine of the contact angle is taken as 2/π [3]. The 

absorbed values increase with the water-cement ratio. A deviation is obvious for 

M12 with blast furnace slag cement and a high cement content. 

The values of r calculated from the sorptivity are always larger than calculated 

from the penetration coefficient. This feature is certainly due to the fact 

that the single size tube model is only a rough approximation of reality. It reality 

the smallest pores have the greatest capillary suction form while the complete 

filling of the pores are lacking behind. This means that the penetration coefficient 

should take into account smaller pores than the sorptivity does. 

As the absolute values of r are concerned these are rather large compared to 

pore sizes which are calculated from many intrusion experiments [8]. Obviously, 

the pores which are reached by the organic fluid are the larger ones and the very 

small pores are either filled by water of are inaccessible due to other reasons, for 

instance due to the viscosity of the fluid or of the size of the molecule. This 

could also mean that the model of the sharp wetting front is questionable. On the 

other hand, it means that the selection of various fluids could give an impression 

of the pore sizes which can be detected. 

CONCLUSIONS 

∗ The experiments with n-decane have proven the capillary suction law which 

states that the absorbed volume and the penetration depth are a function of 

the square root of time. 

∗ The water-cement ratio is the main parameter governing the absorption 

properties. 

∗ The infiltration test with 20 kPa leads only to a minor increase of the penetration 

and absorption. 

75 



∗ The pore sizes determined from the absorption test are larger than from the 

penetration test indicating a different access to pores by different mechanisms. 

∗ Maximum aggregate size and various cement contents lead to different 

physical properties. 

REFERENCES 

[1] Reinhardt, H. W. (ed.): Penetration and permeability of concrete: barriers 

to organic and contaminating liquids. London: E&FN Spon, 1997 

[2] Aufrecht, M.: Beton als sekundäre Dichtbarriere gegenüber umweltgefährdenden 

Flüssigkeiten - Technologie und Konzept für den Schadensfall, 

Dissertation Universität Stuttgart, 1994 

[3] Sosoro, M.: Modell zur Vorhersage des Eindringverhaltens von organischen 

Flüssigkeiten in Beton, DAfStb, H. 446, Berlin 1995 

[4] Brauer, N.: Analyse der Transportmechanismen für wassergefährdende 

Flüssigkeiten in Beton zur Berechnung des Medientransports in ungerissene 

und gerissene Betondruckzonen, DAfStb, H. 524, Berlin 2002 

[5] Paschmann, H., Grube, H., Thielen, G.: Untersuchungen zum Eindringen 

von Flüssigkeiten in Beton sowie zur Verbesserung der Dichtheit des Betons. 

DAfStb, H. 450, Berlin 1995 

[6] DAfStb Guideline "Betonbau beim Umgang mit wassergefährdenden Stoffen", 

Part 4, Berlin 1996 

[7] Pfingstner, A.: Determination of concrete pore structure parameters from 

penetration tests with n-decane, Otto Graf Journal 10 (1999), pp. 113-127 

[8] Reinhardt, H.-W., Gaber, K.: From pore size distribution to an equivalent 

pore size of cement mortar. In: Materials & Structures 23 (1990), pp. 3-15 

76

Analysis of crystalline materials contained in a palestine kohl vessel from the 4th century A.D. 

ANALYSIS OF CRYSTALLINE MATERIALS PRESERVED IN A PAL- 

ESTINE KOHL VESSEL FROM THE 4 TH CENTURY A.D. 

UNTERSUCHUNGEN AM KRISTALLINEN INHALT EINES KA- 

JALGLASES AUS PALÄSTINA, 4. JH. A.D. 

ANALYSE DU CONTENU CRISTALLIN D'UN RECIPIENT A KHOL 

DE PALESTINE DATANT DU 4ÈME SIÈCLE A.D. 

Friedrich Grüner 

SUMMARY 

The crystalline content of a Late Roman glass vessel used to hold cosmetic 

eye shadow (kohl) was analysed. The analytical techniques used were X-ray 

powder diffraction and scanning electron microscopy. The materials detected are 

described, indicating that they may have been used as kohl. 


Es wurde der kristalline Inhalt eines spätrömischen Doppelglasgefäßes aus 

Palästina mit Röntgenpulverdiffraktometrie und am Rastelektronenmikroskop 

untersucht. Der Gefäßinhalt wurde wahrscheinlich als Augenschminke (Kajal) 

benutzt. 

RESUME 

Le contenu cristallin d'un récipient romain provenant de Palestine a été analysé 

au diffractomètre poudre aux rayons X et au microscope électronique à balayage. 

Le contenu du récipient était probablement utilisé comme maquillage 

pour les yeux (khôl). 

KEYWORDS: kohl, glass vessel, galena, anglesite, cerussite, x – ray diffraction 


The following is a report on a study of the materials contained in a double – 

tube flask from the collection of the “Württembergisches Landesmuseum” in 

Stuttgart. A typical glass vessel for holding cosmetic eye – paints might have 

one, two or four individual tubes. 

77 


F. GRÜNER 

Studies of kohl previously reported in the literature have dealt with Egyptian 

material /1/ and Late Roman to Byzantine material /2/. Galena (lead sulfide) 

and the basic copper carbonate, malachite were widely used in Egypt for this 

purpose. Both types were used in the Predynastic period, but the use of malachite 

had stopped by the end of the New Kingdom. The use of galena continued 

into the Coptic period. In Palestine glass vessels from the mid 4 th to early 7 th 

century only galena was found in previous studies /2/. 

The double – tube flask of the collection of the Württembergische Landesmuseum 

was made out of one long glass bleb, which had been divided into two 

sections. Than both sections had been blown separately. The glass is light green 

in colour and shows many bubbles. Both tubes contained a chunk of altered, 

dark grey kohl (Fig. 1). One tube with a broken fragment shows part of a bronze 

or copper rod, sticking in the altered kohl. The total height of the vessel is 9.9 

cm, the diameter of each tube is approximately 1.7 cm. 

Fig. 1: Double tube flask made of light green glass with 4 bails. One tube is broken and 

shows the preserved residue of the kohl and the corroded bronze rod. 

78


Fig 2: Detailed photograph of the rod sticking in the kohl. 

In Fig 2 some details of the chunk and the sticking rod are shown. The rod 

is partly covered with green, blue – green and red coloured corrosion products. 

The surface of the kohl is dark grey in colour and shows sometimes metallic 

brightness. 

2. EXPERIMENTAL PROCEDURES 

For the detailed analyses at least one sample of the altered kohl was removed 

from each tube. The samples were prepared for X – ray diffraction 

(XRD), using a Siemens D 500 diffractometer. Scanning electron microscopy 

with energy dispersive spectrometry (SEM/EDS) were used to identify the 

chemical elements present. A Camscan scanning electron microscope including 

a Noran Voyager energy dispersive x-ray analyzer was used for microscopic investigation. 

3. ANALYTICAL RESULTS 

The results of the analyses of the kohl vessels are presented below. Two 

samples were removed from the surface of the solid chunk in both tubes. In both 

samples the most common alteration products of galena, anglesite (PbSO4) and 

cerussite (PbCO3) were present in major amounts. But galena (PbS) was also 

observed in minor amounts (see Fig. 2). Both samples are nearly identical in 

composition and could not be distinguished with x – ray diffraction. 

79 


F. GRÜNER 

800 

700 

600 

500 

400 

300 

(Counts) 200 

qr 

S 

100 

10 

1 

0 

5 10 20 30 40 50 60 70 

SchminkeP2 - File: SchminkeP2.RAW - Type: 2Th/Th locked - Start: 5.000 ° - End: 70. 

83-1720 (C) - Anglesite - Pb(SO4) - Y: 69.31 % - d x by: 1. - WL: 1.5406 - Orthorhombi 

36-1461 (*) - Anglesite, syn - PbSO4 - Y: 75.98 % - d x by: 1. - WL: 1.5406 - Orthorho 

2-Theta - Scale 

05-0592 (I) - Galena, syn - PbS - Y: 14.79 % - d x by: 1. - WL: 1.5406 - Cubic - a 5.936 

47-1734 (*) - Cerussite, syn - PbCO3 - Y: 10.30 % - d x by: 1. - WL: 1.5406 - Orthorho 

Fig. 2: XRD plot of the powdered kohl sample. Anglesite and cerussite occurred as common 

alteration products, but galena is also present. 

It is reasonable that finely ground galena for use as kohl would have 

enough time to alterate into anglesite and cerrusite during ca. 1500 years of storage 

under unknown archaeological conditions. The analyses of a small piece of 

the rod, sticking inside the kohl showed cuprite and brochantite (see Fig. 3). 

3000 

2000 

1000 

(Counts) 

qr 

S 

600 

500 

400 

300 

200 

100 

10 

0 

5 10 20 30 40 50 60 70 

Schminke P1 - File: SchminkeP1.RAW - Type: 2Th/Th locked - Start: 5.000 ° - End: 70 

75-1531 (C) - Cuprite - Cu2O - Y: 90.16 % - d x by: 1. - WL: 1.5406 - Cubic - a 4.26000 

43-1458 (I) - Brochantite-M - Cu4SO4(OH)6 - Y: 2.72 % - d x by: 1. - WL: 1.5406 - Mon 

2-Theta - Scale 

83-1720 (C) - Anglesite - Pb(SO4) - Y: 3.25 % - d x by: 1. - WL: 1.5406 - Orthorhombic 

36-1461 (*) - Anglesite, syn - PbSO4 - Y: 8.38 % - d x by: 1. - WL: 1.5406 - Orthorhom 

47-1734 (*) - Cerussite, syn - PbCO3 - Y: 0.61 % - d x by: 1. - WL: 1.5406 - Orthorhom 

Fig. 3: XRD plot of a mixed sample with kohl (anglesite, cerussite) and alteration products of 

the bronze rod (cuprite, brochantite). 

80


The complete vessel was placed under the scanning electron microscope 

for further investigations. The original surface of the kohl was studied in the 

broken tube. Elemental analysis showed high concentrations of lead and sulphur 

and some copper in the surrounding material of the rod, indicating a bronze alloy 

or copper metal. Other elements like Sb were absent, eliminating the use of 

stibnite as possible component in the kohl material. 

Fig. 4: Elemental analysis of a galena cube at the surface showing mainly Pb, 

S is buried by the Pb peak. 

Photomicrographs of the surface are shown in Fig. 5 and 6. The kohl consists 

of a very fine grained groundmass with hypidiomorphic intergrown cubes 

of galena. 

Fig. 5: Photomicrograph of the fine grained groundmass with intergrown galena cubes. 

81 


F. GRÜNER 

4. CONCLUSIONS 

Fig. 6: Detailed photomicrograph of a galena cube. 

In this study evidence was found only for galena as material used for the 

production of kohl. Both flasks contained identical materials. For its use as ancient 

make up (eye shadow) it should be ground very fine. It is reasonable to 

assume that most of the galena is altered to anglesite and cerussite during the 

long period of storage under unknown archaeological conditions. 

5. ACKNOWLEDGEMENT 

The author wish to thank Mrs. Dr. Honroth at the Württembergisches Landesmuseum 

for providing the sample material. 

REFERENCES 

[1] LUCAS, A., 1962: Ancient Egyptian Materials and Industries, 4th ed., revised 

by J.R. Harris (Edward Arnold, London, 1962), pp 80-84 

[2] BLANCHARD, W.D., STERN, E.M., STODULSKI, L.P., 1992: Analysis of Materials 

contained in Mid-4 th to Early 7 th Century A.D. Palestinian Kohl Tubes, 

Mat. Res. Soc. Symp. Proc. Vol. 267, pp 239-254 

82

Acoustic emission analysis of SFRC beams under cyclic bending loads 

ACOUSTIC EMISSION ANALYSIS OF SFRC BEAMS UNDER CYCLIC 

BENDING LOADS 

SCHALLEMISSIONSANALYSE AN STAHLFASERBETON UNTER 

ZYKLISCHEN BIEGEVERSUCHEN 

ANALYSE DES ÉMISSIONS ACOUSTIQUES DE BETONS 

RENFORCES PAR FIBRES D'ACIER SOUS FLEXION CYCLIQUE 

Florian Finck 

SUMMARY 

To further understand the failure processes within steel fibre reinforced 

concrete members under cyclic load, a series of 3-point bending tests was 

performed on notched beams using quantitative acoustic emission (AE) 

measurements. AE measurements supplement the mechanical test data by 

providing a large quantity of information about the progress of damage in terms 

of time, location and cause. Quantitative analysis of acoustic signals consists of 

an accurate localization of the fracturing and under certain assumptions an 

inversion for the moment tensor can be performed to gain information about the 

total energy released and the orientation of the rupture plane. After 

decomposition of the moment tensor, the type of rupture process can be 

quantified and visualized using ostensive crack models like those for shear and 

opening. In this article some first results of the fatigue test series and the 

analysis of the AE-data are presented. 


Zur Untersuchung von Schädigungsprozessen innerhalb 

stahlfaserbewehrter Betonbauteile unter zyklischer Last wurde eine Reihe von 

3-Punkt-Biegeversuchen durchgeführt und die auftretenden Schallereignisse 

aufgezeichnet. Neben den mechanischen Prüfdaten können so Informationen 

über den Zeitpunkt und den genauen Ort der fortlaufenden Schädigung 

gewonnen werden. Darüber hinaus kann unter bestimmten Voraussetzungen 

eine Inversion auf den Momententensor durchgeführt werden, welcher 

bruchmechanische Parameter, wie z. B. die Bruchenergie und die Orientierung 

der Bruchflächen enthält. Nach einer geeigneten Zerlegung des Tensors können 

die enthaltenen Bruchmoden quantifiziert und durch anschaulich Bruchmodelle, 

83 


F. FINCK 

wie die des Öffnungs- oder des Scherbruches beschrieben werden. In diesem 

Artikel werden erste Ergebnisse der Ermüdungsversuche und der 

Schallemissionsanalyse vorgestellt. 

RESUME 

Afin d'analyser les processus de détérioration à l'intérieur d'éléments en 

béton armé de fibres d'acier sous chargement cyclique, nous avons réalisé une 

série d'essais de flexion 3 points et enregistré les émissions acoustiques. Ainsi 

nous avons pu gagner, outre les données mécaniques, des informations sur la 

nature, le moment et le lieu exacts des émissions acoustiques, et, par là, sur la 

progression de la rupture. Dans certaines conditions, le tenseur des moments 

peut être calculé par inversion. Celui-ci contient des informations sur l'énergie 

libérée et l'orientation de la surface de rupture. Après la décomposition du 

tenseur de moment, les modes de rupture peuvent être quantifiés et visualisés à 

l'aide de modèles simples, comme ceux pour le cisaillement et l'ouverture. Dans 

cet article les premiers résultats des essais de fatigue et de l'analyse des 

émissions acoustiques sont présentés. 

KEYWORDS: fatigue test, steel fibre reinforced concrete, acoustic emission, 

moment tensor 

INTRODUCTION 

Steel fibre reinforced concrete (SFRC) has been in use since the late 60s, 

mainly as shotcrete for underground constructions and flooring. Some 

advantages of SFRC are a minimization of crack widths and permeability or an 

increased toughness. Although various basic works on the behaviour of SFRC 

members have been published [e.g. WEILER 2000], there remain open questions 

about the mechanical laws and processes that exist during failure. The 

interaction between steel fibre reinforcement and a cementitious matrix, as well 

as the characterization of failure of SFRC members, are mayor topics of the 

subproject A6 in the collaborative research centre SFB 381. 

In a fatigue test series with steel fibre reinforced concrete (SFRC) beams 

under cyclic 3-point bending load we studied the behaviour of ongoing failure. 

Thereby, the investigation of acoustic emissions under changing conditions 

(e. g. load, amplitude and frequency) was the main focus, not an accurate 

statistical investigation of the members. The external, i. e. visible, fatigue is 

84


given by mechanical test data containing the deflection in dependence on load 

and the number of load cycles. Additionally, acoustic emission analysis yields 

information about the internal processes of failure, which correspond to the 

emission of seismic energy due to cracking. Each single crack (event) is 

localized and for a selection of events moment tensors are evaluated. A suitable 

decomposition of this tensor [JOST & HERMANN 1989] yields parameters such as 

the energy released during rupture, the orientation and the size of the rupture 

plane and a combination of ostensive fracture modes. From these parameters the 

stress regime in the member and the mechanics of failure can be derived. 

SETUP OF THE FATIGUE TEST SERIES 

For the test series beams with dimensions 15 cm X 15 cm X 70 cm with a 

1.5 Vol.% reinforcement of Dramix ® RC 80/60 BN steel fibres (length: 60 mm, 

diameter: 0.75 mm) were used. On the bottom surface in the middle of the beam 

a notch with a depth of approximately 3.3 cm caused a well-defined start of a 

crack. The transmission of force by the servo hydraulic 100 kN test frame was 

realized using three steel cylinders. The two fixed lower supports had a distance 

of 60 cm and the upper support at the centre was free to rotate around the 

longitudinal axis of the beam to avoid torsional stress. Figure 1 shows details of 

the test setup, with the AE sensors attached to the specimen. 

Figure 1: Sketch of a notched SFRC beam under a cyclic 3-point bending load. AE sensors 

are mounted around the area of damage. 

85 


F. FINCK 

Piston displacement, load and crack opening were recorded over time. The 

acoustic emissions were recorded by an 8-channel transient recorder with a 

sample rate of 2.5 MHz per channel and an amplitude resolution of 12 Bit. Eight 

piezo-electric accelerometers were evenly distributed on the surfaces of the 

beam. 

First, the range of the failure load was evaluated in two static bending tests. 

Although, a large variation of this value has to be expected due to a change in 

the distribution of fibres and the composition of concrete in the beam, this value 

provides an estimate of the load profile for the dynamic tests. During the 

dynamic tests, load, amplitude and frequency were adapted to the progress of 

damage and the AE activity. 

PRESENTATION OF THE RESULTS 

In the following section the results of one test run are presented. A load of 

7.5 kN was applied statically before the load cycles began. In the second plot in 

figure 2 the different phases with changing load, amplitude and frequency 

during the fatigue test are labelled and coded by blue (dark grey) and green 

(light grey) respectively to be identified in the load over crack opening 

(displacement) plot and the crack opening over time or cycles plot. The 

frequency of the sinusoidal load cycles was 1 Hz from phase G. From that point 

the number of load cycles equals time in seconds plus 5000. On the bottom a 

histogram of the acoustic emission activity can be correlated with the ongoing 

failure in the beam. 

With each increase of the maximum load the crack opening, as well as the 

AE activity, increases rapidly but a relaxation is visible with the continuation of 

the test. The extending areas of the hysteretic ellipses in the load over crack 

opening plots are another indicator for the damage progress. 

86


Figure 2: Mechanical test data of one fatigue test. From top: load deflection curve, 

load over time profile, crack opening over time and the AE activity. 

87 


F. FINCK 

During the test a total of 385 acoustic emissions were recorded from which 

377 could be localized. The data quality was very good regarding noise due to 

the cyclic bending and the accuracy of the localization lies in a range of about 

1 cm. 

Figure 3 shows the located events from three prospectives: from above, a 

front view and a side view. The markers representing the sources of the acoustic 

emissions are given in the legend, corresponding to the test periods starting with 

the according labels (see also figure 2, 2 nd plot). This illustrates the temporal 

growth of the damaged zone. 

Figure 3: Projection of the localization of acoustic emissions. The different markers 

correspond to various test periods, as indicated in the legend. 

88


Nearly all acoustic emissions lie in the central region of the specimen in the 

vicinity of the main crack. Due to the steel fibres, some smearing of the damage 

zone takes place. The early events come from the lower half of the specimen 

since the crack starts in the edge of the notch due to tension. Then steel fibres 

are activated and accommodate load as they are pulled out. The crack grows 

towards the top of the specimen under a relative constant spatial AE activity 

from the complete region under fatigue. The width of the damage zone in ydirection 

is more or less in the range of the fibre length (i. e. 60 mm). This 

suggests that always the short end of the fibre is being pulled out, as expected. 

THE INVERSION OF MOMENT TENSORS 

To gather more information about the mechanical reasons of failure, we 

calculate moment tensors with a relative moment tensor inversion (RMTI) 

technique developed by DAHM 1993. The application and some theory of this 

method on acoustic emission data has been described previously [e. g. FINCK 

2002, FINCK 2001, GROSSE 1999]. An advantage of the RMTI is the elimination 

of the Green’s functions [AKI & RICHARDS 1980] of the medium by an inversion 

for a cluster of events. Two circles in figure 3 indicate the orientation of two 

clusters of 16 events each which were inverted for their moment tensors. C1 is a 

cluster from very early events in the tension zone, events in C2 originate from 

an advanced stage of the test. 

Figure 4: Radiation patterns of seismic energy and results from the moment tensor inversion 

for selected events from cluster C1 an C2 (see figure 3). Mr is the relative seismic moment, 

ISO is the isotropic component of the event and DC is the double-couple portion of the 

deviatoric component. 

89 


F. FINCK 

A combination of two different crack modes is expected for the performed 

test. First, the opening of the main crack should radiate energy similar to event 

EV 29 in cluster C1. An opening mainly perpendicular to the vertical cracksurface 

with particle motion outwards parallel to the y-axis and a significant 

remaining isotropic component. This conforms to mode 1. Second, a great 

number of events should correspond to the pull-out of steel fibres. Mainly a 

double couple mechanism for shear failure is expected in this case, with a small 

isotropic component only (conforming mode 2 or 3). 

A selection of the results is shown in figure 4. The first row contains results 

for cluster C1, the second for C2. Under the top view projection of the radiation 

patterns of seismic energy the relative seismic moment Mr, the isotropic 

component ISO and the double-couple portion DC of the deviatoric component 

are given with errors. The best results from a boot strap analysis [EFRON & 

TIBSHIRANI, 1986] can be found in the brackets. The majority of the moment 

tensors consist of a very small positive isotropic component. For event EV 29 

the errors are very high, so the results must be doubted, though the radiation 

pattern fits to first expectations. For the other events in C1 the DC component is 

rather small. The deviatoric components of these events can not be explained by 

one pure shear crack. Other deviatoric phenomena seem to take place. But the 

early events in C1 vary from the results for C2. The events occurred at an 

advanced stage of the test, where the fibre-pull out seems to be the major reason 

for acoustic activity. Here, the DC component is large. 

The results have a great stability for a changing composition of the 

investigated clusters. In earlier investigations the results for single events were 

dependent on the composure of the cluster, meaning that the existence or nonexistence 

of other events had an influence on the results. Also the errors are 

small. 

CONCLUSIONS 

We successfully obtained high quality acoustic emission data from cyclic 

bending tests of steel fibre reinforced concrete beams. The majority of these 

events could be localized and an inversion for the moment tensor of a selection 

of events was performed. Stable results from the moment tensor inversion can 

partially be correlated with the expected mechanisms of failure – an opening of 

the crack (mode 1) and mainly the pull-out of fibres (mode 2 or 3). Acoustic 

90


emission analysis helps understanding complex mechanisms of failure even over 

a large period of time. 

The decomposition of the moment tensor into crack modes known from 

geological investigations seem not to be suitable for experimental data from the 

laboratory. A decomposition taking crack modes from engineering models in to 

account, is needed. This subject will be of intensive interest in future 

ACKNOWLEDGEMENTS 

These investigations are part of our work in the collaborative research 

centre SFB 381 at the University of Stuttgart which is financially supported by 

the Deutsche Forschungsgemeinschaft (DFG). We gratefully acknowledge this 

support. The author would also like to thank Lindsay Linzer, Rock Engineering 

Dept., CSIR Miningtek for providing the radiation pattern generator. 

REFERENCES 

AKI, K., RICHARDS, P.G.: Quantitative Seismology; Volume 1. Freeman and 

Company, New York, 1980. 

DAHM, T.: Relativmethoden zur Bestimmung der Abstrahlcharakteristik von 

seismischen Quellen. Dissertation, Universität Karlsruhe, 1993. 

EFRON, B. TIBSHIRANI, R.: Bootstrap methods for standard errors, confidence 

intervals and other measures of statistical accuracy. Statistical Science 1, 

pp.54-77, 1986. 

FINCK, F.: Application of the moment tensor inversion in material testing. Otto- 

Graf-Journal, Vol. 12, pp. 145-156, 2001. 

FINCK, F., MOTZ, M., GROSSE, C.U., REINHARDT, H.-W., KRÖPLIN, B.: 

Integrated Interpretation and Visualization of a Pull-Out Test using Finite 

Element Modelling and Quantitative Acoustic Emission Analysis. Online 

publication: http://www.ndt.net/article/v07n09/09/09.htm, 2002. 

GROSSE, C.U.: Grundlagen der Inversion des Momententensors zur Analyse von 

Schallemissionsquellen. Werkstoffe und Werkstoffprüfung im Bauwesen. 

Festschrift zum 60. Geburtstag von Prof. Dr.-Ing. H.-W. Reinhardt, Libri 

BOD, Hamburg, pp. 82-105, 1999. 

91 


F. FINCK 

JOST, M.L., HERMANN, R.B.: A students guide to and review of moment tensors. 

Seism. Res. Letters, Vol. 60, pp. 37-57, 1989. 

WEILER, B.: Zerstörungsfreie Untersuchung von Stahlfaserbeton. Dissertation an 

der Universität Stuttgart, Shaker Verlag, 2000. 

92

About the Improvement of US measurement techniques 

ABOUT THE IMPROVEMENT OF US MEASUREMENT TECHNIQUES 

FOR THE QUALITY CONTROL OF FRESH CONCRETE 

GERÄTETECHNISCHE FORTSCHRITTE BEI DER QUALITÄTS- 

SICHERUNG VON FRISCHBETON MIT ULTRASCHALL 

AMÉLIORATION DES TECHNIQUES DE MESURE ULTRASONIQUES 

POUR LE CONTRÔLE DE QUALITÉ DU BÉTON FRAIS. 

Christian U. Grosse 

ABSTRACT 

Over the last decade a testing method based on ultrasound was developed 

at the Institute of Construction Materials of the University of Stuttgart to control 

the hardening process of cementitious materials by means of non-destructive 

testing. This paper describes the systematic improvement and re-design of the 

testing system and the investigation methods. 

ÜBERSICHT 

Am Institut für Werkstoffe im Bauwesen der Universität Stuttgart wurde in 

den letzten zehn Jahren ein Ultraschallverfahren zur die Analyse des Erstarrens 

und Erhärtens von zementgebundenen Materialien entwickelt. Der Artikel 

beschreibt die fortdauernde Verbesserung der Messtechnik im Hinblick auf die 

Qualitätskontrolle von Frischbeton und –mörtel. 

RESUME 

A l'université de Stuttgart, un procédé ultrasonique de contrôle de la prise 

et du durcissement des matériaux cimentaires a été développé au courant des dix 

dernières années. L'article présent décrit l'amélioration continue des dispositifs 

et de la procédure de mesurage en ce qui concerne le contrôle de la qualité des 

béton et mortiers frais. 

KEYWORDS: Fresh concrete, non-destructive testing, ultrasound 

93 


C. U. GROSSE 

INTRODUCTION 

Nowadays the characterization of cement-based materials during the 

stiffening process by ultrasound measurement techniques is well established. 

This paper deals with the ultrasound technique used in through transmission. In 

numerous publications [e. g. GROSSE & REINHARDT 1994, GROSSE ET AL. 1999, 

REINHARDT ET AL. 1999a] the patented test method [REINHARDT ET AL. 1999b] 

developed at the University of Stuttgart was described earlier. Methods based on 

ultrasound are better suited for the characterization of the setting and hardening 

of cement based materials than traditional test methods like the Vicat-needletest, 

the penetrometer test or the flow test, because the travel time, the 

attenuation and the frequency content of ultrasound waves sent through the 

material are closely correlated with the elastic properties of concrete or mortar. 

These parameters can be continuously monitored during the stiffening giving a 

comprehensive picture instead of snapshots of workability for example. 

A sophisticated device was developed and numerous experiments have 

been conducted in the past, investigating the influence of water-to-cement ratio, 

the type of cement, the use of additives and admixtures, the air bubble content 

and so far, for the setting and hardening of concrete or mortar. Newer features 

are the extraction of the initial and final setting time out of the signals [GROSSE 

& REINHARDT 2000] and the parallel registration of the state of hydration. 

However, the earlier described device lacks of handiness and several features, 

which could improve the art of such measurements further. 

EVOLUTION AND SURVEY OF DEVICES EXISTING AT THE 

UNIVERSITY OF STUTTGART 

The first measurements to control the setting and hardening of concrete at 

the University of Stuttgart using ultrasound are dated back to the early 1990’s. 

These experiments have been conducted in the frame of a research project 

sponsored by the German Reinforced Concrete Committee (DAfStb, V 345) and 

are published in the 1994 th volume of the Otto-Graf-Journal [GROSSE & 

REINHARDT 1994]. The tests were carried out with a rough set-up using a 

container made of 40 mm thick styrene foam (Styropor) plates and the dimensions 

300 mm × 300 mm × 80 mm (Fig. 1). The emitter was a simple steel ball impactor 

dropping a ball of 4 mm diameter on to a small aluminium plate, which was placed 

in contact with the fresh concrete. 

94


Fig. 1: Set-up with steel ball impactor and receiver. 

Dimensions of the container: 300 mm x 300 mm x 80 mm 

Later on the ideas were proofed by numerous students during their Diploma 

thesis, technician and student research assistants. Jochen FISCHER [1994], Bernd 

Weiler and the author [Grosse 1996] developed a device using three long styrene 

foam walls and two smaller rigid side walls out of aluminium plates and the same 

simple impactor as used earlier (Fig. 2). 

Fig. 2: Set-up of the smaller styrene foam container with two aluminium side walls. 

Dimensions of the container: 200 mm x 80 mm x 60 mm. 

95 


C. U. GROSSE 

While the first set-up caused problems to determine the correct travel 

distance of the pulse to the receiver, what is essential for velocity measurements, 

the second set-up was unsatisfactory as well, because of interfering waves 

resulting from the walls. 

A re-design of the device described in REINHARDT ET AL. [1996], WINDISCH 

[1996], HERB [1996] and REINHARDT ET AL. [1998] for concrete measurements 

was patented later [Reinhardt et al. 1999] and consisted of a mould completely 

out of PMMA of the dimensions 160 mm × 200 mm × 70 mm (Fig. 3), but the 

handling of this device was poor and the leakiness of the container caused a 

penetration of fluids especially during the compaction process. However, the 

device was modified by BEUTEL [1999] and tested to be suitable for field 

measurements. 

Fig. 3: Set-up of the first container out of PMMA only. 

Dimensions of the container: 160 mm x 200 mm x 70 mm. 

In the meantime the development of a test set-up adjusted to mortar 

materials run parallel. Due to smaller grain sizes (usually less 2 mm) the 

dimensions of a mortar device can significantly be reduced. Not all steps of the 

development can be described in detail. Figure 4 gives an impression of the 

iterative process of finding a suitable shape for the mould. The final container 

[GROSSE ET AL. 1999] had two walls of PMMA and a U-shaped rubber foam 

with an inner volume of 40 cm³ for the mortar (Fig. 5). 

96


Fig. 4: Evolution of containers, tested for mortar applications. 

There are two main advantages in respect to the concrete set-up. The 

amount of material necessary to be tested is significantly reduced and so is the 

amount of waste. Secondly, the pulse is not excited by an impactor, what is an 

advantage in terms of reliability and handiness. 

Fig. 5: Final set-up of the mortar device showing the mould (rubber foam and PMMA-walls) 

and the transducers. 

97 


C. U. GROSSE 

Consequently, a new device, illustrated in Fig. 6, was developed by 

STEGMAIER [2000] and Herb, whereby the dimensions were changed to 400 mm 

× 59 mm ×130 mm in accordance to the smaller mortar device. Similar to 

former concrete devices the wave is generated using a steel ball exciter, referred 

to as Ultrasound Impactor (USIP), hitting a small plate fixed on the PMMA 

casing. The resulting excitation can be seen as broad banded, having a relatively 

wide frequency bandwidth of up to 100 Hz. 

Fig.6: FreshCon device for concrete measurements developed on the basis of the older mortar 

device (see Fig.5). 

Though many difficulties were eliminated the system still shows up 

unresolved problems. Specifically, a wave travelling through the container wall 

which onset is detected before the irradiating primary wave can be observed. 

Further on, the energy evolution during the hardening of concrete is still difficult 

to analyze since the steel ball transmitter USIP, as a mechanical system, 

provides unreliable energy data and the plate where the steel ball is shot on 

easily disbond so that the coupling of the excited energy into the PMMA 

container changes during tests. These factors influence the obtained results and 

the reproducibility of tests. 

98


To summarize the pros and cons of the concrete device the following 

statements can be given: 

• Less reproducibility of impact energy results in energy determination uncertainties. 

• Contact problems of steel plate at PMMA container (delaminations). 

• Unreliable generation of impacts due to steel balls sticking in the impactor rod. 

• Possible side wall waves disturbing the measurement at early ages during 

investigations of very “slow” materials. 

• Pressure air equipment necessary for the impactor. 

It should be stated, that at the end of 2000 no possibility to record the 

hydration temperature in the same sample during the ultrasound measurements 

as a secondary control technique was available using the existing FreshCon 

software. 

ANALYSIS METHODS 

Using ultrasound methods the degree of hardening is characterized by the 

change of significant parameters. Not only the travel time of the ultrasonic pulse 

through the testing device, consequently the velocity of compressional waves 

but also the frequency content and the relative energy are recorded. 

On the basis of suitable parameters, e.g. the frequency content of the signal 

over the time, additionally a wavelet transformation (WT) is carried out in order 

to gather as much information as possible from the raw signal to evaluate 

concrete and mortar, respectively. The program AutoCWT, able to apply the 

WT was implemented by MANOCCHIO [2001], where the calculation kernel is 

taken from the program IWB-CWT, coded by BAHR [2001a]. More information 

about the application of wavelets in the characterization of the setting and 

hardening of cementitious materials can be obtained from Grosse [2001], 

GROSSE & REINHARDT [2001] or MANOCCHIO [2001]. 

Further on as a new feature of the FreshCon system the ability to record the 

temperature evolution over the time is introduced as well as the determination of 

the associated hydration heat, following DIN-EN 196 part 9. 

99 


C. U. GROSSE 

Fig. 7: Set-up (top) for measurements of elastic parameters (velocity, energy, frequency) as 

well as the temperatures. Bottom: Screenshot of the new program version 2.04 of FreshCon. 

100


MEASUREMENTS OF HYDRATION TEMPERATURES 

The program FreshCon was extended to enable temperature measurements 

using the multi-channel National instrument computer board NI 4351. A 

screenshot of this program version is represented in Fig. 7, where also a picture 

of the test setup is given. The temperature distribution occurring during the 

hydration process is a characteristic for the state of hardening of cement-based 

materials. Therefore, statements can be deduced according the relations between 

two different materials. It should be mentioned that the hydration process in the 

semi adiabatic container in comparison to the testing device for ultrasound 

measurements is faster due to the accumulation of heat in the temperature 

container. Consequently, the sound velocity and temperature distribution cannot 

be correlated directly. A picture of the testing device for the determination of the 

heat of hydration, taken from KÖBLE [1999] is given in Fig. 8. 

328 

181 

232 

292 

cable Thermokabel to digital geführt thermometer in 

einem Plexiglasröhrchen 

wooden Holzdeckel lid 

rubber Zellgummidichtung, foam, d = 20 d mm = 2cm 

seal Dichtung 

polystyrene, Polystyrol, d = d = 5cm 50 mm 

seal Dichtung 

plate Plexiglasscheibe of lucite to fix zum thermocouple fixieren 

des Temperaturfühlers 

can Mörteldose, filled with h mortar = 12cm 

Dewar Dewar-Gefäß container Ø 16cm 160 mm 

polystyrene Polystyrolscheibe 

air Luft 

polystyrene 

Stützvorrichtung 

to hold 

aus 

Dewar 

Polystyrol 

container in upright position 

wooden Gehäuse box aus Holz 

Fig. 8: Set-up of the calorimeter device according KÖBLE [1999], dimensions in mm. 

Regarding the determination of the heat of hydration DIN EN 196 - 9 is 

followed, accordingly. The aim of the semi-adiabatic method, namely the 

Langavant - method, applicable to mortar, is the determination of the released 

amount of heat during the hydration process. For this purpose the online version 

of the program FreshCon, implemented by BAHR [2001b], was modified. The 

system is now able to record the temperature in the calorimeter (Fig. 8), the 

temperature in the tested material (Fig. 7, top) and the air temperature. All these 

data are obtained automatically and stored together with the data of the 

ultrasound measurements. A typical result is represented in Fig. 9, showing all 

101 


C. U. GROSSE 

three temperatures as a function of the concrete age. The temperature effect is 

dominant at the curve obtained using the calorimeter (straight line) due to the 

semi-adiabatic conditions in the Dewar container. Testing concrete materials a 

hydration effect is clearly seen at the temperature data obtained in the ultrasound 

container (dotted line) compared to the air temperature (dashed line). 

Air and Concrete temperature [°C] 

45 

40 

35 

30 

25 

20 

15 

10 

5 

Hydration 

Concrete 

Air 

Rilem Round Robin Tests 

iBMB: 18/19 Apr 2002 

Mixture: RB03 (concrete) 

Remarks: very dry, water added 

0 

0 

0 200 400 600 800 1000 1200 1400 

Time [min] 

Fig. 9: Results of temperature measurements using the new FreshCon version. 

DEVELOPMENT OF A NEW CONCRETE FRESHCON DEVICE 

Comparing the two devices for mortar and concrete measurements, the 

advantages of the mortar set-up should be summarized: 

• Good reproducibility of signal generation (energy). 

• Easy onset time determination due to signals with good reproducibility. 

• No pressure air equipment necessary. 

• Full automatic measurement and storing of waveforms. 

• Automatic determination of velocity and energy as well as additional parameters. 

• Full control of measurement parameters. 

To enhance the handling of concrete experiments accordingly, a new 

design of the device shown in Fig. 6 was suggested. For the new device the 

102 

45 

40 

35 

30 

25 

20 

15 

10 

5 

Hydration Heat [°C]


impactor was replaced by an US transmitter in combination with a wideband 

power amplifier and a function generator. No pressure air is need for this device; 

a control sensor next to the impactor recording the emitting pulse is no longer 

required. 

velocity [m/s] 

2000 

1500 

1000 

500 

0 

velocity impactor 

velocity piezo 

rel. energy impactor 

rel. energy piezo 

0 120 240 360 480 600 720 

age [min] 

Fig. 10: Comparison experiments between impactor and US emitter. 

In several preliminary experiments the new set-up was tested by 

MANOCCHIO [2001] to compare the results of measurements by the impact 

generated signals and by piezo-electric emitters in parallel (Fig. 10). The two 

curves at the bottom of the right side in Fig. 10, recorded at the same time using 

the same material, represent the velocity evaluation of gypsum. Gypsum was 

used as a test material due to its fast hydration evolution. The two curves at the 

top position in Fig. 10 represent the relative energy. A decrease of the velocity 

and energies values is caused by shrinkage effects. Both curve pairs look very 

similar in respect to differently used pulse generation methods. 

This successful first test triggered the re-design of the concrete device (as 

well as of the mortar device). A flow chart of the new experimental set-up is 

given in Fig. 11. The electronic pulse is generated by a frequency generator and 

amplified by a power amplifier. Broadband piezo-electric transducers generate 

the ultrasound signal to be transmitted through the material. A transducer of the 

same type is used as a receiver and the signal is passed through a pre-amplifier 

to the PC-board A/D-converter, denoted as “computer-based signal processing” 

in Fig. 11. Special attention is given to the correct trigger time of the signal, 

what is essential for velocity measurements. A power amplifier of the companies 

KROHN-HITE CO. or DEVELOGIC GMBH is used along with sensors of the 

company VALLEN INSTRUMENTS. 

103 

1 

0.1 

0.01 

1E-3 

1E-4 

1E-5 

1E-6 

Otto-Graf-Journal Vol. 13, 2002 

rel. energy [-]

C. U. GROSSE 

Testing device 

(concrete or mortar) 

Fig. 11: Flow chart of newly developed FreshCon experiments. 

The new container/sensor design for concrete as well as for mortar 

experiments is shown in Fig. 12, demonstrating the similarity of these two. The 

U-shaped rubber in the middle of the container is essential. Regarding the 

concrete device, a special “long wall” container was produced for very “slow” 

materials to avoid waves propagating along the walls to be faster than the direct 

waves. The distance of the screw joints can be adjusted to the material 

properties. 

Fig. 12: Re-designed FreshCon container/sensor for mortar (left) and concrete (right) 

measurements. 

First experiments in the frame of a master thesis [KALCKBRENNER 2002] 

and during round robin test of a RILEM technical committee showed very 

promising results. 

104

ROUND ROBIN TESTS – STATUS 


The International Union of Testing and Research Laboratories for Materials 

and Structures, RILEM, as a non profit-making, non-governmental technical 

association is structured in groups of international experts, the so called 

Technical Committees (TC). In the framework of advanced testing of cementbased 

materials during setting and hardening the TC 185 - ATC organized a 

round robin test series. The purpose of these tests is to assess the capability of 

existing test methods based on non-destructive techniques in terms of suitability, 

sensitivity and accuracy. Results will be summarized in a state of the art report 

and a test recommendation is planned to be released. In the context of providing 

a direct comparability, experiments are carried out by different members at the 

same place using the same charge of materials/mixtures. 

The technical realization of the experiments is in the responsibility of the 

TC secretary (C. Grosse) and the local organizers. The ongoing test series 

started in 2001 with experiments in Vaulx-en-Velin (France) and was continued 

in Evanston/Chicago (USA) in spring 2002 and Brunswick (Germany) in 

summer 2002. The next round robin test is scheduled for spring 2003 in Delft 

(The Netherlands). In detail the following groups have been involved so far: 

• Ecole Nationale des Travaux Publics de l’Etat (ENTPE), Vaulx-en-Velin, 

France; Dr. L. Arnaud and Prof. C. Boutin. 

• Center for Advanced Cement-Based Materials (ACBM) at Northwestern 

University, Illinois, USA; Prof. S. Shah and Dipl.-Ing. T. Voigt. 

• Institute of Structural Materials, Solid Structures and Fire Protection 

(iBMB) of the Technical University of Brunswick, Germany; Prof. H. 

Budelmann, Dipl.-Math. M. Krauß. 

• Fraunhofer Institute for Non-Destructive Testing (IZFP) in Saarbrücken, 

Germany; Dr. G. Dobmann and Dr. B. Wolter. 

• Institute of Construction Materials (IWB) at the University of Stuttgart, 

Germany; Prof. H.-W. Reinhardt, Dr. C. Grosse and Dipl.-Ing. A. Kalckbrenner 

(M.Sc.). 

An experimental test program was compiled to be the basis for all 

experiments [GROSSE & REINHARDT 2002]. Six different mixtures are 

recommended to be tested – five other mixtures are tested additionally. Some of 

the results obtained by the Institute of Construction Materials (IWB) at the 

105 


C. U. GROSSE 

University of Stuttgart are published by KALCKBRENNER [2002] and correlated 

to the results of other groups. A comprehensive report will follow. 

To give an example of the data obtained during one test series Fig. 13 

demonstrate the variation of the velocities over the age of the material. 

Concerning these velocities an S-shaped curve is typical for cementitious 

materials. After a certain time at the beginning, while the velocity variation is 

small, the gradient is increasing significantly. Regarding the data RE5 from a 

mix with added retarder this increase occurs relatively late. To make the basic 

statements more evident the curves are smoothed and bad data points are 

removed. It is obvious that concrete mixes are “faster” than mortar mixes in 

respect to hardening, while the RE5 mix with retarder is the “slowest“ material. 

Velocity [m/s] 

5000 

4000 

3000 

2000 

1000 

0 

smoothed! 

RE1: concrete, w.c. 0.45 

RE2: concrete, w.c. 0.60 

RE3: mortar, w.c. 0.60 

RE4: mortar, with plasticizer 

RE5: mortar, with retarder 

RE6: mortar, with air entrainer 

0 200 400 600 800 1000 1200 1400 

Age [min] 

Fig. 13: Comparison of the velocity measurements testing mixtures RE 1-6. 

It should be stressed that only material properties related to the elastic behavior 

can be analyzed with ultrasound techniques. As far as the chemical properties 

are not related to the elastic properties, other measurement techniques have to be 

used in combination with ultrasound to get more data. The results of the round 

robin tests should indicate the value of the described ultrasound throughtransmission 

technique in comparison to other techniques like ultrasound 

reflection, nuclear magnetic resonance, electric and maturity methods. 

106

SUMMARY AND OUTLOOK 


The measuring device developed at the University of Stuttgart is able to 

analyze the setting and hardening of cementitious materials in a comprehensive 

way. The method is based on ultrasound and can be used for numerous 

applications, where reliable and reproducible data are required, what addresses 

material parameters like the water-to-cement-ratio, the type of cement or the 

effect of additives as retarders or accelerators. At the concreting site, where 

efficiency and a low budget are boundary conditions, the application of this new 

technique can help to enhance the stability during construction or the progress of 

the construction work saving both: time and money. Some examples are the 

development of admixtures, the in-situ quality control, the slip form concreting 

or the precasting. Certainly, the applications are not restricted to cementitious 

materials. 

Further improvements are concerning the velocity evaluation. Since the 

device consist of an analogue-to-digital converter of 5 MHz only, the resolution 

of the velocity calculations varies over age. Actually, the resolution decreases 

with increasing velocities. This is the reason of the so-called bit-pattern 

occurring usually at ages of 400 minutes and later. To ease the interpretation the 

velocity curves are smoothed using adjacent averaging (10 points), but it is 

suggested to plot the original data points into the smoothed curves as well. 

Using the offline version of the FreshCon picking algorithm the data can be reevaluated 

after the test concerning the onset times of the signals only. 

Surprisingly, curves re-picked by the operator are usually very similar to the 

automatically processed data so that a time consuming manually picking is not 

improving the results anymore. 

Formerly, the comparison of energy evaluation results was sophisticated 

due to the application of two different devices. Energy values as measured by 

the FreshCon software are basing on the squared amplitudes of the signal 

beginning at the signals onset of compressional waves. These values strongly 

depend on the energy released by the impact to the container. The 

reproducibility of the transmitter energy is low of impactor devices compared to 

devices using an ultrasound emitter. Changing the set-up as described made the 

interpretations regarding energies more reliable. There is still the disadvantage 

of energies emitted by piezo-driven devices to be of several magnitudes lower 

than impactor pulses. A new impactor device without pressure-air giving broadband 

pulses of reproducible magnitude is under development. 

107 


C. U. GROSSE 

Talking about the scientific aspects of the ultrasound technique, the method 

developed at the University of Stuttgart is under further progress. This is 

especially true concerning wavelet algorithms. The degree of automatization is 

enhanced and additional analysis techniques will be implemented in future. 

With regard to the international activities of the RILEM technical 

committee more information can be obtained from the author or at the TC’s 

homepage: http://www.rilem.org/atc.html. Colleagues working in this scientific 

field are offered to collaborate in this initiative. 


The described design and re-design of ultrasound devices are the result of 

many years of scientific work. It is difficult to address the thanks to everybody 

who was involved. However, some colleagues should be mentioned in no 

particular order: Dr. B. Weiler, Dipl.-Ing. J. Fischer, Dipl.-Ing. I. Kolb, Dipl.- 

Ing. N. Windisch, Dipl.-Ing. A. Herb, Dipl.-Ing. S. Köble, Dipl.-Ing. R. Beutel, 

Dipl.-Ing. C. Manocchio, Dipl.-Ing. A. Kalckbrenner (M.Sc.), Mr. G. Bahr and 

Mr. G. Schmidt. A special acknowledgement is going to Prof. H.-W. Reinhardt 

who initiated this research project and contributed during the years in numerous 

ways. 

The results shown regarding measurements in the frame of the RILEM TC 

185-ATC were obtained during a collaboration with the research group of Dr. 

Laurent Arnaud, Laboratoire Géomatériaux, Département Génie Civil et 

Bâtiment, of the Ecole Nationale des Travaux Publics de l’Etat (ENTPE) in 

Vaulx-en-Velin near Lyon, France. 

108

REFERENCES 


Bahr, G.: Entwicklung von Algorithmen für die kontinuierliche Wavelet Transformation 

mit LabView. University of Stuttgart, internal report (2001a). 

Bahr, G.: Bedienungsanleitung FreshCon 2.04. University of Stuttgart, Institute 

of Construction Materials, manual (2001b). 

Beutel, R.: Praktische Anwendbarkeit der Ultraschallwellenmessung als Instrument 

zur Bestimmung des Erhärtungsgrades von Beton. Diploma thesis, 

University of Stuttgart, 2000. 

Fischer, J.: US-Messungen an Frischbeton. Diploma thesis, University of 

Stuttgart, 1994. 

Grosse, C. U., H.-W. Reinhardt: Continuous ultrasound measurements during 

setting and hardening of concrete. Otto-Graf-Journal 5 (1994), pp 76-98. 

Grosse, C. U.: Quantitative zerstörungsfreie Prüfung von Baustoffen mittels 

Schallemissionsanalyse und Ultraschall. PhD Thesis, University of Stuttgart, 

1996, 168 pages. 

Grosse, C. U., B. Weiler, A. Herb, G. Schmidt, K. Höfler: Advances in ultrasonic 

testing of cementitious materials. Festschrift zum 60. Geb. von Prof. 

Reinhardt (C. U. Grosse, Ed.), Libri publishing company, Hamburg (1999), 

pp. 106-116. 

Grosse, C. U., H.-W. Reinhardt: Ultrasound technique for quality control of 

cementitious materials. Proc. of 15. World Conf. on NDT, Rom 2000, (on 

CD-ROM and in the internet at www.ndt.net). 

Grosse, C. U.: Verbesserung der Qualitätssicherung von Frischbeton mit 

Ultraschall. Concrete Plant and Precast Technology, Vol. 67, No. 1 (2001), 

pp. 102-104. 

Grosse, C. U., H.-W. Reinhardt: Fresh concrete monitored by ultrasound 

methods. Otto-Graf-Journal Vol. 12 (2001), pp. 157-168. 

Herb, A.: Frischbeton: Korrelation zwischen Ergebnissen klassischer Konsistenzmessungen 

und Ultraschall-Verfahren. Diploma thesis, University of 

Stuttgart, 1996. 

Kalckbrenner, A.: On the modification of non-destructive ultrasound 

measurement techniques for quality control of cement based materials. 

Master Thesis, University of Stuttgart, 2002. 

109 


C. U. GROSSE 

Köble, S.: Physikalisch-chemischer Hintergrund des Hydratationsvorgangs von 

Frischmörtel im Hinblick auf Ultraschalluntersuchungen. Diploma thesis, 


Manocchio, C.: Verwendung der Wavelet-Transformation zur Charakterisierung 

von Frischbeton mittels Ultraschall. Diploma thesis, University of Stuttgart, 

2001. 

Reinhardt, H.-W., C. U. Grosse: Setting and hardening of concrete continuously 

monitored by elastic waves. Proc. of the Int. RILEM Conf. "Prod. methods 

and workability of concrete", Paisley/Schottland (1996), pp. 415-425. 

Reinhardt, H.-W., C. U. Grosse, A. Herb: Kontinuierliche Ultraschallmessung 

während des Erstarrens und Erhärtens von Beton als Werkzeug des 

Qualitätsmanagements. Deutscher Ausschuss für Stahlbeton, No. 490 

(1999a), pp. 21-64. 

Reinhardt, H.-W., C. U. Grosse, A. Herb, B. Weiler, G. Schmidt: Verfahren zur 

Untersuchung eines erstarrenden und/oder erhärtenden Werkstoffs mittels 

Ultraschall. Patent pending under No. 198 56 259.4 at the German Patent 

Institution, Munich (1999b). 

Stegmaier, M.: Zerstörungsfreie Prüfung des Erstarrens und Erhärtens von 

Beton – Weiterentwicklung des Ultraschallprüfverfahrens. Diploma thesis, 


Windisch, N.: Untersuchung der Erhärtung von Beton – hochfester Beton bzw. 

Fließbeton – mit Ultraschallwellen, Diploma thesis, University of Stuttgart, 

1996. 

110

Α δισχρετε βονδ µοδελ φορ 3∆ αναλψσισ οφ τεξτιλε ρεινφορχεδ ανδ πρεστρεσσεδ χονχρετε ελεµεντσ 

A DISCRETE BOND MODEL FOR 3D ANALYSIS OF TEXTILE REIN- 

FORCED AND PRESTRESSED CONCRETE ELEMENTS 

DISKRETES VERBUNDMODELL FÜR 3D-FE-BERECHNUNGEN VON 

TEXTILBEWEHRTEN UND VORGESPANNTEN BETONKONSTRUK- 

TIONEN 

UN MODELE DISCRET DE L'ADHERENCE POUR L'ANALYSE 3D DE 

STRUCTURES EN BETON RENFORCEES ET PRECONTRAINTES 

AVEC DES ARMATURES TEXTILES 

Μαρκυσ Κρ⎫γερ, ϑο�κο Ο�βολτ, Ηανσ−Ω. Ρεινηαρδτ 

SUMMARY 

Τεξτιλε ρεινφορχεδ χονχρετε στρυχτυρεσ σηοω σεϖεραλ σιγνιφιχαντ αδϖανταγεσ 

χοµπαρεδ το στεελ ρεινφορχεδ χονχρετε στρυχτυρεσ ωηιχη αρε ωελλ κνοων υπ το 

νοω. Ηοωεϖερ σοµε δισαδϖανταγεσ λικε τηε λοω υτιλιζατιον φαχτορ οφ τηε τεξτιλε 

ρεινφορχεδ ελεµεντσ βεχοµε οβϖιουσ. Ασ ιν ανψ ρεινφορχεδ στρυχτυρε, α τρανσφερ 

οφ φορχεσ φροµ ρεινφορχεµεντ το χονχρετε ισ αχχοµπλισηεδ τηρουγη βονδ. Τηερε− 

φορε υνδερστανδινγ ανδ φυρτηερ ιµπροϖεµεντ οφ βονδ προπερτιεσ βετωεεν τεξτιλε 

ανδ χονχρετε ισ ιµπορταντ. Ιν τηε παπερ βονδ προπερτιεσ βετωεεν διφφερεντ τεξτιλεσ 

ανδ ηιγη περφορµανχε φινε γραιν χονχρετε αρε δισχυσσεδ. 

Νυµεριχαλ σιµυλατιονσ ωιτη α ϖαριατιον οφ ινπυτ δατα ωερε περφορµεδ υσινγ 

α νονλινεαρ φινιτε ελεµεντ χοδε βασεδ ον τηε µιχροπλανε µοδελ φορ χονχρετε ανδ 

τηε δισχρετε βονδ µοδελ. Τηε βονδ µοδελ ισ βασεδ ον α δισχρετε Φινιτε ελεµεντσ 

φορµυλατιον ωηιχη χαν βε υσεδ φορ στεελ ρεινφορχεδ χονχρετε ασ ωελλ. Τηε νυ− 

µεριχαλ σιµυλατιονσ ανδ ϖαριατιον οφ παραµετερσ σηοω τηε ινφλυενχε οφ διφφερεντ 

βονδ χηαραχτεριστιχσ οφ τεξτιλε ρεινφορχεµεντσ ανδ τηερεφορε γιϖε σοµε ηιντσ ον 

ποσσιβλε οπτιµισατιον οφ τεξτιλε στρυχτυρεσ. 


Ωιε βερειτσ ιν δερ νευερεν Λιτερατυρ ερωηντ, ζειγεν τεξτιλε Βεωεηρυνγσ− 

µατεριαλιεν ιν ϖερσχηιεδενεν Ανωενδυνγσγεβιετεν δευτλιχηε ςορτειλε γεγεν⎫βερ 

κονϖεντιονελλερ Σταηλβεωεηρυνγ. Αβερ αυχη εινιγε Ναχητειλε ωιε διε γερινγε 

νυτζβαρε Φεστιγκειτ τεξτιλερ Βεωεηρυνγ ιν Βετονβαυτειλεν υνδ διε δαµιτ ϖερβυν− 

δενεν ηοηεν Κοστεν σπρεχηεν γεγεν εινεν Εινσατζ σολχηερ Βεωεηρυνγσµατερια− 

111 

Οττο−Γραφ−ϑουρναλ ςολ. 13, 2002

Μ. ΚΡ⇐ΓΕΡ, ϑ. Ο�ΒΟΛΤ, Η.−Ω. ΡΕΙΝΗΑΡ∆Τ 

λιεν. Ιµ Αλλγεµεινεν κοµµτ δεµ ςερβυνδ ζωισχηεν Βεωεηρυνγ υνδ Βετον βει 

δεραρτιγεν ςερβυνδωερκστοφφεν εινε ηοηε Βεδευτυνγ ζυ, σινδ διεσε δοχη υντερ 

ανδερεµ µα⇓γεβενδ φ⎫ρ δασ Τραγϖερηαλτεν. Ιµ ϖορλιεγενδεν Βειτραγ ωερδεν 

δαηερ ωεσεντλιχηε ςερβυνδειγενσχηαφτεν ϖερσχηιεδενερ τεξτιλερ Βεωεηρυνγεν ιν 

Βετον δισκυτιερτ υνδ ερλυτερτ. 

Ανηανδ ϖον νιχητλινεαρεν Φινιτε−Ελεµεντ−Βερεχηνυνγεν µιτ Παραµετερϖα− 

ριατιονεν ωιρδ ειν νευεσ ςερβυνδµοδελλ ζυρ Χηαρακτερισιερυνγ τεξτιλερ Βεωεη− 

ρυνγεν ιν Βετον ϖοργεστελλτ. ∆ασ ιν δεν ΦΕ−Χοδε ΜΑΣΑ εινγεβυνδενε ςερ− 

βυνδµοδελλ βασιερτ ιµ Ωεσεντλιχηεν αυφ δεν γλειχηεν Ανναηµεν ωιε σιε φ⎫ρ δεν 

Σταηλ−/Βετονϖερβυνδ γελτεν υνδ ωυρδε ιν εινιγεν ωενιγεν Πυνκτεν φ⎫ρ τεξτιλε 

Βεωεηρυνγεν ανγεπασστ. Νυµερισχηε Σιµυλατιονεν ζειγεν, ωιε Εινφλ⎫σσε τεξτι− 

λερ Βεωεηρυνγεν αυφγρυνδ υντερσχηιεδλιχηερ Στρυκτυρ υνδ Αρτ βερ⎫χκσιχητιγτ υνδ 

ωιε ζυδεµ τεξτιλε Βεωεηρυνγεν ηινσιχητλιχη δεσ Τραγϖερηαλτενσ τεξτιλβεωεηρτερ 

Βαυτειλε οπτιµιερτ ωερδεν κ⎞ννεν. 

RESUME 

Λεσ αρµατυρεσ τεξτιλεσ οντ πλυσιευρσ αϖανταγεσ σιγνιφιχατιφσ παρ ραππορτ αυξ 

αρµατυρεσ χονϖεντιοννελλεσ εν αχιερ. Χεπενδαντ χερταινσ ινχονϖ⎡νιεντσ χοµµε 

λε βασ ταυξ δ∋εξπλοιτατιον δε λα ρ⎡σιστανχε δε λ∋αρµατυρε τεξτιλε ετ λεσ χο⎦τσ ⎡λεϖ⎡σ 

θυι εν ρ⎡συλτεντ σ∋οπποσεντ ◊ λευρ αππλιχατιον ◊ γρανδε ⎡χηελλε. Λ∋αδη⎡ρενχε εντρε 

λ∋αρµατυρε ετ λε β⎡τον ϕουε υν ρ⎮λε ιµπορταντ δανσ λεσ µατ⎡ριαυξ χοµποσιτεσ, ελλε 

εστ σουϖεντ δ⎡χισιϖε πουρ λε χοµπορτεµεντ σουσ χηαργε δ∋υνε στρυχτυρε. ∆ανσ 

λ∋αρτιχλε πρ⎡σεντ, λεσ χαραχτ⎡ριστιθυεσ δε λ∋αδη⎡ρενχε δε διφφ⎡ρεντεσ αρµατυρεσ τεξ− 

τιλεσ σοντ δ⎡χριτεσ ετ δισχυτ⎡εσ. 

∆εσ σιµυλατιονσ νυµ⎡ριθυεσ αϖεχ υνε ϖαριατιον δεσ παραµ⎝τρεσ οντ ⎡τ⎡ εφ− 

φεχτυ⎡σ εν υτιλισαντ δεσ ⎡λ⎡µεντσ φινισ νον−λιν⎡αιρεσ βασ⎡σ συρ λε µοδ⎝λε ∀µιχρο− 

πλανε∀ πουρ λε β⎡τον ετ λε νουϖεαυ µοδ⎝λε δισχρετ δε λ∋αδη⎡ρενχε. Λε µοδ⎝λε δε 

λ∋αδη⎡ρενχε εστ βασ⎡ συρ υν µοδ⎝λε δισχρετ δε λ∋αδη⎡ρενχε θυι πευτ ⎢τρε ⎡γαλεµεντ 

εµπλοψ⎡ πουρ λε β⎡τον αϖεχ υνε αρµατυρε εν αχιερ. Λεσ σιµυλατιονσ νυµ⎡ριθυεσ 

ετ λα ϖαριατιον δεσ παραµ⎝τρεσ µοντρεντ λ∋ινφλυενχε δε διφφ⎡ρεντεσ χαραχτ⎡ριστιθυεσ 

δε λ∋αρµατυρε τεξτιλε. Ον πευτ εν δ⎡δυιρε δεσ µεσυρεσ πουρ οπτιµισερ λε χοµπορ− 

τεµεντ δεσ στρυχτυρεσ αϖεχ δεσ αρµατυρεσ τεξτιλεσ. 

ΚΕΨΩΟΡ∆Σ: υνχοατεδ τεξτιλεσ, ιµπρεγνατεδ τεξτιλεσ, χονχρετε, Χαρβον, ΑΡ 

γλασσ, βονδ, βονδ µοδελ, 3∆ ΦΕ αναλψσισ, πρεστρεσσ 

112

INTRODUCTION 


Βονδ βεηαϖιουρ οφ τεξτιλε ρεινφορχεµεντ ιν χονχρετε ισ εξπεχτεδ το ϖαρψ 

φροµ τηατ οφ ΦΡΠ βαρσ ορ χονϖεντιοναλ στεελσ βαρσ. Μοστ τεξτιλε ροϖινγσ υσεδ ασ 

χονχρετε ρεινφορχεµεντ χονσιστ οφ τηουσανδσ οφ σινγλε φιλαµεντσ ανδ τηερεφορε 

χαν νοτ βε δεφινεδ ασ α σινγλε ροδ. Ιφ συχη α ροϖινγ ισ εµβεδδεδ ιν χονχρετε τηε 

σηαπε οφ τηε χροσσ σεχτιον δετερµινεσ τηε βονδεδ αρεα ανδ ιτ µυστ βε χλαριφιεδ 

ηοω µανψ φιλαµεντσ ωερε ιν διρεχτ χονταχτ ωιτη χονχρετε. Α γρεατ δεαλ οφ ρε− 

σεαρχη ηασ βεεν δονε ρεχεντλψ το χηαραχτεριζε βονδ βεηαϖιουρ οφ συχη µυλτιφιλα− 

µεντ ελεµεντσ ιν χονχρετε βυτ θυιτε νεω ιννοϖατιονσ νεχεσσιτατε φυρτηερ ρεσεαρχη 

/ΒΡΑΜΕΣΗΥΒΕΡ, 2000/, /ΝΑΜΜΥΡ, 1989/, /ΟΗΝΟ, 1994/. Μορεοϖερ θυιτε α νυµ− 

βερ οφ εξπεριµενταλ ινϖεστιγατιονσ ηαϖε βεεν χαρριεδ ουτ το υνδερστανδ βονδ βε− 

ηαϖιουρ οφ πρεστρεσσεδ ανδ/ορ ιµπρεγνατεδ τεξτιλεσ ορ ροϖινγσ. 

Ονε παραµετερ τηατ µαψ στρονγλψ ινφλυενχε τηε βονδ περφορµανχε ισ τηε διφ− 

φερενχε ιν τηε χοεφφιχιεντ οφ τηερµαλ εξπανσιον φροµ τηατ οφ στεελ ορ χονχρετε. Ιτ 

ισ αλσο κνοων τηατ τρανσϖερσε πρεσσυρε ιµπροϖεσ βονδ ωηιχη ισ νεγλεχτεδ ιν 

µανψ βονδ µοδελσ. Ηοωεϖερ, τηισ εφφεχτ σεεµσ το βε νοτ ιµπορταντ φορ εµβεδ− 

δεδ µυλτι−φιλαµεντ ροϖινγσ ωηιχη ηαϖε νοτ βεεν φυλλψ ινφιλτρατεδ ωιτη χεµεντ 

δυε το ϖοιδσ βετωεεν τηε ιννερ φιλαµεντσ. ∆εσπιτε τηισ τηε Ποισσον�σ εφφεχτ βε− 

χοµεσ σιγνιφιχαντ ανδ ινφλυενχεσ τηε τρανσϖερσε στρεσσ φιελδ ιφ τηε ροϖινγ ισ ιµ− 

πρεγνατεδ ανδ/ορ πρεστρεσσεδ. Σοµε τεστ ρεσυλτσ οφ χαρβον ρεινφορχεδ ανδ 

πρεστρεσσεδ σπεχιµεν αρε ιλλυστρατεδ ιν Φιγυρε 1 /ΚΡ⇐ΓΕΡ, 2001Β/. 

P/(c-∆s*), N/mm 

60 

55 

50 

45 

40 

35 

30 

25 

20 

15 

10 

5 

0 

0 1 2 3 

Carbon 

no prestressing 

prestress 150 N/roving 


Carbon, epoxy impreg. 

no prestressing 



4 

slip ∆s*, mm 

Φιγυρε 1: Βονδ στρεσσ περ υνιτ λενγτη ϖερσυσ σλιπ βασεδ ον 20µµ δουβλε σιδεδ πυλλ−ουτ 

(στορεδ ατ 20°Χ, 65% ΡΗ φορ 40 δαψσ) /ΚΡ⇐ΓΕΡ, 2001Β/ 

113 



Ιτ χαν βε σεεν τηατ αν ιµπρεγνατιον οφ α χαρβον ροϖινγ ωιτη αν εποξψ ρεσιν 

γενεραλλψ ρεσυλτσ ιν α βεττερ βονδ ωηερεασ α ροϖινγ τηατ ωασ νοτ ιµπρεγνατεδ 

σηοωσ α λοω µαξιµυµ βονδ στρεσσ ανδ αφτερ βονδ φαιλυρε α ϖερψ λοω φριχτιοναλ 

ρεσιστανχε. Ιτ ισ ασσυµεδ τηατ τηε µαιν ρεασον φορ τηισ ισ τηε ριββεδ συρφαχε 

φορµεδ βψ τηε βινδερ τηρεαδσ ανδ τηε χηανγε οφ τηε ροϖινγ διαµετερ οϖερ ιτσ 

λενγτη, εσπεχιαλλψ ατ τηε χροσσινγ ποιντσ ωηερε τηε περπενδιχυλαρ ωοοφ ροϖινγ ισ 

φιξεδ. Τηε βινδερ τηρεαδσ αρε χαυσεδ βψ τηε ωαρπ κνιττινγ προχεσσ (Φιγυρε 2) ανδ 

ωερε φιξεδ βψ τηε εποξψ ρεσιν. Ιτ χαν βε σεεν φροµ φιγυρε 1 τηατ πρεστρεσσινγ 

λεαδσ το α ηιγηερ βονδ στρενγτη. Ασ δισχυσσεδ αβοϖε, βονδ περφορµανχε οφ τεξτιλε 

ρεινφορχεµεντ ιν χονχρετε δεπενδσ ον µανψ διφφερεντ παραµετερσ. Τηισ λεαδσ υσ 

το χονσιδερ α φορµυλατιον οφ α συχη βονδ µοδελ ιν ωηιχη τηεσε ασπεχτσ ωουλδ βε 

αχχουντεδ φορ. 

10 mm 

Φιγυρε 2: ∆εταιλ οφ αν εποξψ ιµπρεγνατεδ χαρβον φαβριχ 

DISCRETE BOND MODEL FOR FINITE ELEMENT ANALYSIS 

Φορ νυµεριχαλ στυδιεσ τηε βονδ προπερτιεσ βετωεεν τεξτιλεσ ανδ χονχρετε, δισχρετε 

ελεµεντσ ωερε υσεδ. Τηε βονδ µοδελ προποσεδ βψ /Ο�ΒΟΛΤ, 2002/ ηασ τηερεφορε 

βεεν µοδιφιεδ φορ τεξτιλε ρεινφορχεµεντ ανδ υσεδ τογετηερ ωιτη σολιδ φινιτε ελε− 

µεντσ ιν α 3∆ ΦΕ στυδιεσ. 

Ιν τηε νυµεριχαλ στυδιεσ βονδ βετωεεν τηε τεξτιλεσ ανδ χονχρετε ωασ σιµυλατεδ 

βψ δισχρετε βονδ ελεµεντ τηατ ηαϖε ρεχεντλψ βεεν ιµπλεµεντεδ ιντο 3∆ ΦΕ χοδε 

ΜΑΣΑ /Ο�ΒΟΛΤ, 2002/. Χονχρετε, ωηιχη ισ δισχρετιζεδ βψ τηε τηρεε διµενσιοναλ 

φινιτε ελεµεντσ, ισ µοδελλεδ βψ τηε µιχροπλανε µοδελ /Ο�ΒΟΛΤ, 2001/. Τηε βονδ 

ελεµεντσ χοννεχτ τηε χονχρετε φινιτε ελεµεντσ ωιτη τηε ρεινφορχεµεντ τηατ ισ ρεπ− 

ρεσεντεδ βψ τηε τρυσσ φινιτε ελεµεντσ (σεε Φιγυρε 3). Ονλψ δεγρεεσ οφ φρεεδοµ ιν 

τηε βαρ διρεχτιον αρε χονσιδερεδ. Ηοωεϖερ, βεσιδε τηε τανγεντιαλ στρεσσεσ παραλλελ 

το τηε βαρ διρεχτιον, τηε ραδιαλ στρεσσεσ περπενδιχυλαρ το τηε βαρ διρεχτιον αρε 

γενερατεδ ασ ωελλ. Ιτ ισ ασσυµεδ τηατ ατ α γιϖεν σλιπ τηε ραδιαλ στρεσσ δεπενδσ ον 

114


ερατεδ ασ ωελλ. Ιτ ισ ασσυµεδ τηατ ατ α γιϖεν σλιπ τηε ραδιαλ στρεσσ δεπενδσ ον τηε 

γεοµετρψ οφ τηε βαρ ανδ τηε βαρ στραιν ασ ωελλ ασ ον τηε γεοµετρψ ανδ τηε βουνδ− 

αρψ χονδιτιονσ οφ τηε χονχρετε σπεχιµεν. Τηε ιντεραχτιον βετωεεν τανγεντιαλ ανδ 

ραδιαλ στρεσσεσ ισ αχχουντεδ φορ ιν τηρεε διφφερεντ ωαψσ: (ι) διρεχτλψ, τηε σηεαρ 

στρεσσ δεπενδσ ον τηε νονλοχαλ (ρεπρεσεντατιϖε) ραδιαλ στρεσσ οβταινεδ φροµ τηε 

χονχρετε ελεµεντσ χλοσε το τηε ρεινφορχινγ βαρ, (ιι) τηε λοχαλ στραιν οφ τηε βαρ ελε− 

µεντ ανδ ιτσ λατεραλ εξπανσιον ορ εξτενσιον ανδ (ιιι) ινδιρεχτλψ, ιν α ωαψ τηατ τηε 

λαργερ σηεαρ στρεσσ (ηιγηερ βονδ στρενγτη δυε το λαργερ ριβσ ορ ρουγηνεσσ οφ τηε 

βαρ ελεµεντ) χαυσε ηιγηερ αχτιϖατιον οφ στρεσσεσ ιν τηε ραδιαλ διρεχτιον. 

Ιν τηε πρεσεντ µοδελ, σπλιττινγ οφ χονχρετε ισ ινδιρεχτλψ αχχουντεδ φορ. 

Ναµελψ τηε ιντεραχτιον βετωεεν σηεαρ ανδ ραδιαλ στρεσσεσ ρεσυλτσ ιν χορρεσπονδ− 

ινγ τανγεντιαλ τενσιλε στρεσσεσ τηατ χαυσεσ χραχκινγ οφ τηε συρρουνδινγ νον−λινεαρ 

χονχρετε ελεµεντσ ανδ, τηερεφορε, φαιλυρε οφ βονδ ρεσιστανχε. 

Concrete 

element 

Bond element 

(zero width) 

fibre element 

Repeated 

nodes 

Φιγυρε 3: Βονδ ελεµεντσ ωιτη ζερο ωιδτη. 

Bond stress-slip relation in a 2D consideration 

Τηε εξπεριµενταλ εϖιδενχε /ΧΕΒ ΒΥΛΛΕΤΙΝ 230, 1996/ ινδιχατεσ τηατ τηε 

λοαδ τρανσφερ βετωεεν ρεινφορχεµεντ ανδ χονχρετε ισ αχχοµπλισηεδ τηρουγη βεαρ− 

ινγ οφ τηε ρεινφορχεδ στεελ λυγσ ον συρρουνδινγ χονχρετε ανδ τηρουγη φριχτιον. Ασ 

δισχυσσεδ βψ /ΨΑΝΚΕΛΕςΣΚΨ, 1987/, τηε τοταλ βονδ ρεσιστανχε χαν βε δεχοµ− 

ποσεδ ιντο τωο χοµπονεντσ: (ι) µεχηανιχαλ ιντεραχτιον χοµπονεντ !µ, ανδ (ιι) 

φριχτιον χοµπονεντ !φ. Τηε φριχτιον χοµπονεντ χαν βε σεπαρατεδ ιντο α ρεσιδυαλ 

φριχτιον !ρ ανδ α ϖιργιν φριχτιον !ϖ χοµπονεντ. Τηε ρεσιδυαλ φριχτιον ρεπρεσεντσ 

φριχτιοναλ ρεσιστανχε υπον σλιπ ρεϖερσαλ ωηερεασ τηε ϖιργιν φριχτιον χοµπονεντ ισ 

δυε το τηε αδδιτιοναλ φριχτιοναλ ρεσιστανχε δεϖελοπεδ υπον λοαδινγ το πρεϖιουσλψ 

υνδεϖελοπεδ σλιπ λεϖελσ. Ιτ ισ ασσυµεδ τηατ τεξτιλε ρεινφορχεµεντ βεηαϖεσ σιµι− 

λαρλψ ασ στεελ ρεινφορχεµεντ δοεσ, ωιτη τηε διφφερενχε τηατ µαινλψ τηε αδηεσιον οφ 

τηε τεξτιλε ανδ τηε ρουγηνεσσ οφ τηε συρφαχε ιµπροϖε τηε µεχηανιχαλ βονδ ινστεαδ 

οφ τηε στεελ λυγσ. 

115 



Βασεδ ον τηε εξπεριµενταλ ρεσυλτσ /ΕΛΙΓΕΗΑΥΣΕΝ, 1983/, /ΜΑΛςΑΡ, 1992/ 

ανδ ασ ωελλ δοχυµεντεδ βψ /ΛΟΩΕΣ, 2002/, τηε βονδ σλιπ ρελατιονσηιπ οφ στεελ 

ρεινφορχεµεντ ιν χονχρετε χαν βε δεσχριβεδ βψ τηε παραµετερσ τηατ αρε συµµα− 

ριζεδ ιν Ταβλε 1. Τηε σαµε παραµετερσ αρε υσεδ φορ τεξτιλε ρεινφορχεµεντ, βυτ ιν α 

σλιγητλψ διφφερεντ µαννερ. Τηε χυρϖε οφ τηε βονδ στρεσσ ϖερσυσ σλιπ ρελατιονσηιπ 

υσεδ φορ τηε νυµεριχαλ στυδιεσ ισ ιλλυστρατεδ ιν Φιγυρε 4. 

Ταβλε1: Συµµαρψ οφ τηε µοδελ παραµετερσ 

∆εσχριπτιον οφ τηε µοδελ παραµετερ Μοδελ παραµετερ 

πεακ µεχηανιχαλ βονδ στρενγτη !µ = !µ,0 ∀ ∗ [ΜΠα] 

πεακ φριχτιοναλ βονδ στρενγτη !φ = !φ,0 ∀ ∗ [ΜΠα] 

πεακ ϖιργιν φριχτιον βονδ στρενγτη !φ,ϖ = (1−0.4) !φ [ΜΠα] 

πεακ ρεσιδυαλ φριχτιον βονδ στρενγτη !φ,ρ = 0.4 !φ [ΜΠα] 

σεχαντ το βονδ ρεσπονσε χυρϖε φορ ινιτιαλ λοαδινγ κσεχ [ΜΠα/µµ] 

σλιπ ατ ωηιχη πεακ βονδ στρενγτη ισ αχηιεϖεδ σ1 = (!µ+!φ)/κσεχ [µµ] 

σλιπ ατ ωηιχη βονδ στρενγτη βεγινσ το δεχρεασε σ2 = σ1+σ2∗ [µµ] 

σλιπ ατ ωηιχη µεχηανιχαλ βονδ ρεσιστανχε ισ λοστ σ3 [µµ] 

τανγεντ το τηε λοαδ−δισπλαχεµεντ χυρϖε υπον υνλοαδινγ κυνλοαδ [ΜΠα/µµ] 

ινιτιαλ τανγεντ το τηε βονδ−σλιπ ρεσπονσε κ1 [ΜΠα/µµ] 

τανγεντ το τηε βονδ−σλιπ χυρϖε ατ πεακ ρεσιστανχε κ2 = α⋅κσεχ [ΜΠα/µµ] 

∗ ∀ σεε νεξτ χηαπτερ 

Τηε παραµετερ !µ,0 ανδ !φ,0 ρεπρεσεντ τηε στρενγτη οφ τηε µεχηανιχαλ ανδ φριχτιοναλ 

χοµπονεντ (συβσχριπτ µ ανδ φ), ρεσπεχτιϖελψ, φορ τηε χασε οφ νο χονφινινγ πρεσ− 

συρε, νο δαµαγε ανδ ελαστιχαλλψ βεηαϖεδ ρεινφορχινγ βαρ ελεµεντ. 

Υπ το τηε σλιπ σ1 ατ ωηιχη πεακ βονδ στρενγτη ισ ρεαχηεδ (σεε Φιγυρε 4), αλλ 

ρεσπονσε χυρϖεσ αρε δεφινεδ βψ Μενεγοττο−Πιντο (ΜΠ) εθυατιον /ΜΕΝΕΓΟΤΤΟ− 

ΠΙΝΤΟ, 1973/. Τηε χυρϖε δεφινεσ α χυρϖε χοννεχτινγ τωο λινε σεγµεντσ ανδ ιτ 

ρεαδσ: 

1 

! ∀ 

# 1 ∃Ρ 

τ () σ = τ⋅τ ! 0 = σ! ⋅ % β + (1−β) ⋅∋ & ⋅τ 

Ρ ( 

% ) 1+ σ! 

∗ & 

+ , 

ωηερε β ισ τηε ρατιο βετωεεν τηε ταργετ ανδ ινιτιαλ τανγεντσ, τ! ανδ σ ! αρε νορµαλ− 

ιζεδ στρεσσ ανδ δισπλαχεµεντ, ρεσπεχτιϖελψ, ανδ Ρ δεφινεσ τηε ραδιυσ οφ τηε χυρϖα− 

τυρε. ανδ σ αρε τηε παραµετερσ το χαλχυλατε τηε αβσολυτε στρεσσ ανδ δισπλαχε− 

τ0 0 

µεντ φροµ τηε νορµαλιζεδ παραµετερσ. 

0 

116 

(1)

β 

κ 

κ 


2 = (2) 

1 

=α⋅ κ2κ σεχαντ , ωιτη 0≤α≤1 

(3) 

σ 

σ! 

= 

σ 

(4) 

Bond 

stress 

0 

(κ −κ 

) (1 ) 

σ σ σ κ 

=⋅ =⋅⋅ −α⋅ 

0 1 

σεχαντ 

κ1−κ2 2 

1 σεχαντ 

κ1− κσεχαντ 

σ κ 

0 0 1 

(6) τ=⋅ 

= m+ f 

0 

m 

f,v 

f,r 

f,r 

s0 

k 1 

k secant 

Cyclic loading 

k 2=α·k secant 

kunload 

Monotonic 

loading 

s1 s3 Slip s 

s2 

Φιγυρε 4: Βονδ στρεσσ−σλιπ ρελατιον οφ τηε βονδ ελεµεντ µοδελ 

Variation of bond strength in a 3D stress field 

(5) 

f= f,r+ f,v 

Α φαχτορ ∀ (σεε Ταβλε 1) αχχουντσ φορ τηε δεπενδενχψ οφ τηε βονδ στρεσσ ον 

τηε στρεσσ−στραιν στατε οφ χονχρετε ανδ στεελ ιν τηε ϖιχινιτψ οφ τηε βονδ ζονε. Ασ α 

ρεσυλτ, τηε τωο διµενσιοναλ βονδ στρεσσ ϖερσυσ σλιπ ρελατιονσηιπ σηοων βεφορε ισ 

ινφλυενχεδ βψ λατεραλ εξπανσιον ορ εξτενσιον ιν διφφερεντ ωαψσ ανδ βεχοµεσ α 

τηρεε διµενσιοναλ µοδελ. 

117 



Τηε παραµετερ ∀ ισ χαλχυλατεδ ασ σηοων ιν εθυατιον (7). Τηρεε παραµετερσ 

αρε χονσιδερεδ: ∀Σ χοντρολσ τηε ινφλυενχε οφ τηε ψιελδινγ οφ στεελ ρεινφορχεµεντ 

ον τηε βονδ ρεσπονσε ανδ ισ σετ το ∀Σ=1 φορ τεξτιλε ρεινφορχεµεντ; ∀Χ αχχουντσ 

φορ τηε ινφλυενχε οφ τηε λατεραλ στρεσσεσ βετωεεν ρεινφορχεµεντ ανδ χονχρετε 

χαυσεδ βψ τηε στρεσσ ιν χονχρετε ανδ τηε λοχαλ στραιν οφ τηε βαρ ελεµεντ ανδ ιτσ 

λατεραλ εξπανσιον ορ εξτενσιον; ∀χψχ χοντρολσ τηε ινφλυενχε οφ τηε λοαδινγ− 

υνλοαδινγ−ρελοαδινγ ον τηε βονδ ρεσπονσε. 

Ω=Ω⋅Ω⋅Ω Σ Χ χψχ 

Ασ σηοων ιν εθυατιον (8) ανδ Φιγυρε 5, τηε παραµετερ ∀Χ, ωηιχη χαν τηεο− 

ρετιχαλλψ ϖαρψ βετωεεν 0 ανδ 2, αχχουντσ φορ τωο διφφερεντ εφφεχτσ. Τηε φιρστ ισ τηε 

ινφλυενχε οφ τηε λατεραλ στραιν οφ τηε στρεσσεδ βαρ ελεµεντ. Τηε παραµετερ ηΡ ισ α 

χονσταντ τηατ ρεπρεσεντσ τηε συρφαχε ρουγηνεσσ οφ τηε ρεινφορχεµεντ βαρ. Χοµ− 

παρεδ το τηε ριββεδ ρεινφορχεµεντ, ηΡ ισ χλοσε ρελατεδ το τηε ηειγητ οφ τηε στεελ 

λυγσ. ισ τηε ρεινφορχεµεντ στραιν, δ = 2ρσ τηε βαρ διαµετερ ανδ µ σ ισ τηε Ποισ− 

σον�σ ρατιο οφ τηε υσεδ ρεινφορχεµεντ ελεµεντ. Τηε φαχτορ αρ χοντρολσ τηε ινφλυ− 

ενχε οφ τηε ραδιαλ χονχρετε στρεσσ ανδ φορ τηε χαλχυλατιονσ ισ σετ το 1. Τηε παραµε− 

σ 

τερ αφ χοντρολσ τηε ινφλυενχε οφ τηε ρουγηνεσσ οφ τηε ρεινφορχεµεντ ηΡ ον τηε 

βονδ ρεσπονσε. Ιν τηε πρεσεντ στυδψ ιτ ωασ σετ το 2. 

σε 

Τηε παραµετερ επ,0 ισ τηε στραιν δυε το πρεστρεσσινγ οφ ρεινφορχεµεντ. Χονσε− 

θυεντλψ ιν τηε χασε οφ πρεστρεσσινγ ανδ νον εξτερναλ λοαδινγ τηε βονδ ισ ιν− 

χρεασεδ ονλψ βψ τηε ραδιαλ στρεσσ ιν χονχρετε νεαρβψ τηε ρεινφορχινγ βαρ. 

# ∃ 

∋ ( 

Ρ 

1 

Ω χ = 1, 0 + τανη ∋ 

ρ φ σ ( σ π,0) ( 

2 

∋ 0,1 φχ 

ρ ( σ 

∋ 1− 

2 

∋ ( 

( ρσ ηΡ) 

( 

) + ∗ ⋅ α⋅−α⋅µ⋅ε−ε⋅ σ 

Ω 

χ 

2,0 

1,0 

0,0 

-3,0 -2,0 -1,0 0,0 1,0 2,0 3,0 

, ( ) 

Ρ 

φ σ σ π,0 2 

ρ 

χ 

σ 1− 

⋅ 

( ρσ + ηΡ) 

Φιγυρε 5: ∆εφινιτιον οφ !Χ ασ α φυνχτιον οφ λατεραλ στρεσσ ανδ στραιν σ−α⋅µ⋅ε−ε⋅ 

118 

0,1 φ 

1 

2 

(7) 

(8)


Τηε ινφλυενχε οφ τηε ραδιαλ στρεσσ ιν χονχρετε ιν τηε ϖιχινιτψ οφ τηε ρεινφορχ− 

ινγ βαρ ισ αχχουντεδ φορ βψ αν αϖεραγε ραδιαλ στρεσσ σ ρ περπενδιχυλαρ το τηε βαρ 

διρεχτιον. Τηε παραµετερ φχ ισ τηε υνιαξιαλ χοµπρεσσιϖε στρενγτη οφ χονχρετε. Ιν 

τηε φινιτε ελεµεντ αναλψσισ τηε αϖεραγε ραδιαλ στρεσσ ασσοχιατεδ το τηε ν−τη βαρ 

ελεµεντ ισ χαλχυλατεδ ασ: 

1 

σ = σ ∆ = − ∆ 

(9) 

Ν Ν 

ι 

ρ − ρ ςι ωιτη ςΡ 

ι 

ς Ρ ι= 1 ι= 1 

ς 

ωηερε ∆ ςι δενοτεσ τηε ϖολυµε ωηιχη χορρεσπονδσ το τηε ι−τη ιντεγρατιον 

ι 

ποιντ οφ τηε φινιτε ελεµεντ ανδ σ τηε στρεσσ περπενδιχυλαρ το τηε ρεινφορχεµεντ. 

ρ 

Ν ισ α τοταλ νυµβερ οφ ιντεγρατιον ποιντσ τηατ φαλλ ιντο α χψλινδερ οφ α διαµετερ ∆ 

(σεε Φιγυρε 6). Ιν τηε πρεσεντεδ µοδελ ∆ ισ ασσυµεδ το βε τηρεε τιµεσ α βαρ δι− 

αµετερ (∆ ≈ 3 δσ). 

Ιν (9) ςΡ ισ τηε ρεπρεσεντατιϖε ϖολυµε, ι.ε. τηε ϖολυµε οφ τηε 

χονχρετε χψλινδερ οφ διαµετερ ∆ τηατ ισ ασσοχιατεδ το τηε τρυσσ φινιτε ελεµεντ 

ωηιχη ρεπρεσεντσ α ρεινφορχινγ βαρ. 

Φιγυρε 6: Ρεπρεσεντατιϖε ϖολυµε 

Εξπεριµεντσ σηοω τηατ φορ χψχλινγ λοαδινγ−υνλοαδινγ−ρελοαδινγ τηε βονδ 

στρενγτη σιγνιφιχαντλψ δεχρεασεσ ωιτη ινχρεασε οφ νυµβερ οφ λοαδινγ χψχλεσ 

/ΕΛΙΓΕΗΑΥΣΕΝ, 1983/, /ΒΑΛΑΖΣ, 1991/. Ιν τηε πρεσεντ µοδελ τηισ εφφεχτ ισ αχ− 

χουντεδ φορ βψ τηε φαχτορ ∀χψχ τηατ ρεαδσ: 

1.1 

# # Λ ∃ ∃ 

εξπ ∋ 1.2 ( 

) ) ∗ ( 

∗ 

Ω χψχ = − ⋅ 

∋ 

∋ ( 

Λ 0 

(10) 

ωηερε # ισ τηε αχχυµυλατεδ σηεαρ ενεργψ δισσιπατιον ανδ #0 ισ α χονσταντ 

ρεπρεσεντινγ τηε αρεα υνδερ τηε µονοτονιχ βονδ−σλιπ χυρϖε οφ ρεσπεχτιϖε σηεαρ 

χοµπονεντ. Τηε αβοϖε εθυατιον ηασ βεεν προποσεδ βψ /ΕΛΙΓΕΗΑΥΣΕΝ, 1983/ ανδ 

ιτ ισ βασεδ ον α λαργε νυµβερ οφ χψχλιχ τεστ δατα οφ στεελ ρεινφορχεδ χονχρετε. 

119 



Σιµιλαρ βεηαϖιορ σεεµσ αλσο το βε αππροξιµατελψ ϖαλιδ φορ τεξτιλε ρεινφορχεδ χον− 

χρετε, ηοωεϖερ, ιτ ισ νοτ χλαριφιεδ υπ το νοω. 

NUMERICAL STUDIES 

Ασ σηοων ιν τηε λαστ χηαπτερ τηε βονδ περφορµανχε οφ τηε µοδελ ισ ινφλυ− 

ενχεδ βψ ∀Χ ωηιχη αχχουντσ φορ: (ι) τηε ινφλυενχε οφ ραδιαλ στρεσσ οβταινεδ φροµ 

τηε συρρουνδινγ χονχρετε ελεµεντσ ανδ (ιι) τηε ινφλυενχε οφ ρεινφορχεµεντ στραιν. 

Το δεµονστρατε τηε εφφεχτσ οφ τηεσε τωο διφφερεντ ινφλυενχεσ νυµεριχαλ στυδιεσ 

ηαϖε βεεν χαρριεδ ουτ. 

Τωο ΦΕ µοδελσ (ΜΙ ανδ ΜΙΙ) ωερε εµπλοψεδ το σηοω τηε ινφλυενχε οφ τηε 

τρανσϖερσε στρεσσ φιελδ ανδ τηε ρεινφορχεµεντ στραιν ον τηε βονδ προπερτιεσ. Τηε 

ΦΕ µεση οφ τηεσε µοδελσ ισ σηοων ιν Φιγυρε 7. Ιτ ρεπρεσεντσ χονχρετε σπεχιµεν 

χονφινεδ ιν διρεχτιον ξ ανδ ψ. Τηε βουνδαρψ χονδιτιονσ ωερε σλιγητλψ διφφερεντ. 

Μοδελ Ι (ΜΙ) ηασ σοµε ρεστραινεδ νοδεσ ιν τηε ζ−διρεχτιον ονλψ ατ τηε βοττοµ συρ− 

φαχε, ωηερε τηε λοαδ ισ αππλιεδ το τηε βαρ ελεµεντ. Ιν Μοδελ ΙΙ (ΜΙΙ) αλλ τηε νοδεσ 

οϖερ τηε σπεχιµεν ηειγητ ωερε φιξεδ ιν τηε ζ−διρεχτιον. Χονσεθυεντλψ, τηε τρανσ− 

ϖερσε στρεσσ φιελδ αρουνδ τηε βαρ ελεµεντ ισ διφφερεντ. Τηε βαρ ελεµεντ ιτσελφ ισ 

πλαχεδ ιν τηε µιδδλε οφ τηε σπεχιµεν ανδ ισ πυλλεδ ιν ζ διρεχτιον ατ τηε ποιντ 

ζ = 20 µµ (βοττοµ συρφαχε). 

Φιγυρε 7: Φινιτε ελεµεντ µοδελσ ωιτη διφφερεντ βουνδαρψ χονδιτιονσ 

Ασ σηοων ιν ταβλε 2, τηε µαιν βονδ παραµετερσ ωερε σετ το χονσταντ φορ αλλ 

χαλχυλατιονσ. Νοτε τηατ τηεσε παραµετερσ ωερε χηοσεν ονλψ το θυαλιτατιϖελψ σηοω 

τηε περφορµανχε οφ τηε µοδελ ανδ ωερε νοτ χαλιβρατεδ φορ τηε υσε ιν τηε πραχτιχαλ 

120


αππλιχατιονσ. Αδδιτιοναλλψ, ιτ ηασ το βε νοτεδ τηατ τηε ρεινφορχεµεντ ισ ασσυµεδ 

το βε λινεαρ ελαστιχ φορ αλλ τηε χαλχυλατιονσ, ι.ε. ∀Σ = 1. 

Ταβλε2: Συµµαρψ οφ τηε υσεδ µοδελ παραµετερσ φορ τηε ρεινφορχεµεντ 

∆εσχριπτιον οφ τηε µοδελ παραµετερ Μοδελ παραµετερ 

πεακ µεχηανιχαλ βονδ στρενγτη !µ = 9.0 [ΜΠα] 

πεακ φριχτιοναλ βονδ στρενγτη !φ = 4.0 [ΜΠα] 

σεχαντ το βονδ ρεσπονσε χυρϖε φορ ινιτιαλ λοαδινγ κσεχ = 50.0 [ΜΠα/µµ] 

σλιπ ατ ωηιχη βονδ στρενγτη βεγινσ το δεχρεασε σ2∗ = 0.03 [µµ] 

σλιπ ατ ωηιχη µεχηανιχαλ βονδ ρεσιστανχε ισ λοστ σ3 = 0.50 [µµ] 

τανγεντ το τηε λοαδ−δισπλαχεµεντ χυρϖε υπον υνλοαδινγ κυνλοαδ = 220.0 [ΜΠα/µµ] 

ινιτιαλ τανγεντ το τηε βονδ−σλιπ ρεσπονσε κ1 = 220.0 [ΜΠα/µµ] 

τανγεντ το τηε βονδ−σλιπ χυρϖε ατ πεακ ρεσιστανχε κ2 = 22.0 [ΜΠα/µµ] 

ραδιυσ οφ τηε χυρϖατυρε Ρ = 8.0 [−] 

Ποισσον�σ ρατιο µσ = 0.5 [−] 

Ψουνγ µοδυλυσ Εσ = 74000.0 [ΜΠα] 

ρεινφορχεµεντ αρεα Ασ = 0.93 [µµ″] 

βαρ διαµετερ 2 ⋅ 

ρσ = 1.0 [µµ] 

συρφαχε ρουγηνεσσ ηΡ = 0.01 [µµ] 

Μορεοϖερ, το δεµονστρατε τηε εφφεχτ οφ πρεστρεσσινγ, τηρεε διφφερεντ χασεσ 

(α,β,χ) ωερε χονσιδερεδ ωιτη: 

(α) Ω = 1.0 

(11) 

(β) 

(χ) 

Χ,α 

Ω = + ∋⋅ 

Χ,β 1.0 Ρ τανη ( 

0,1 φχ 

∃ 

) ∗ 

# ∃ 

∋ ( 

1 

Ω = 1.0 + τανη ∋ 

( ) 

( 

−α⋅µ⋅ε−ε⋅ σ 

Ρ 

Χ,χ φ σ σ π,0 2 

0,1 φχ 

ρ ( σ 

∋ 1− 

2 

∋ ( 

( ρσ + ηΡ) 

( 

∋⋅ 

) ∗ 

#σ 

Influence of the 3D stress field on the bond 

(12) 

(13) 

Ιν Φιγυρε 8 τηε ρεσυλτσ οφ τηε πυλλ ουτ στρεσσ ϖερσυσ σλιπ αρε σηοων. Ιν τηε 

χασε (α) τηε στρεσσ ϖερσυσ σλιπ χυρϖε οφ βοτη µοδελσ λοοκσ αλµοστ τηε σαµε φορ 

πρεστρεσσεδ ανδ νον πρεστρεσσεδ στατε, ι.ε. τηε χονχρετε στραιν ιν ζ διρεχτιον ισ 

νεγλιγιβλε. Τηερεφορε ιτ ισ σηοων ονλψ ονε χυρϖε. Ηοωεϖερ, ιτ χαν βε σεεν τηατ ιφ 

121 



τηε ρεινφορχεµεντ ισ πρεστρεσσεδ µαξιµυµ πυλλουτ στρεσσ ινχρεασεσ σλιγητλψ δυε 

το τηε µορε ηοµογενεουσ βονδ στρεσσ διστριβυτιον οϖερ τηε εµβεδµεντ λενγτη. 

Φορ διφφερεντ χασεσ ατ µαξιµυµ λοαδ αλσο σεε Φιγυρε 9. Τηισ εφφεχτ χαν αλσο βε 

σεεν ιν αλλ τηε χαλχυλατιον δισχυσσεδ λατερ. 

load, N 

1200 

1000 

800 

600 

400 

200 

0 

0,0 0,1 0,2 0,3 0,4 0,5 0,6 

MI & MII, case (a) 

MI & MII, case (a), prestressed 

MI, case (b) 

MI, case (b), prestressed 

MII, case (b) 

MII, case (b), prestressed 

slip, mm 

Φιγυρε 8: Χαλχυλατεδ πυλλ ουτ λοαδ ϖερσυσ σλιπ 

Ιφ χασε (β) ισ χονσιδερεδ ανδ τηε ινφλυενχε οφ τηε ραδιαλ στρεσσ οφ τηε χον− 

χρετε ισ τακεν ιντο αχχουντ, τηε χηανγε οφ βονδ ρεσπονσε βεχοµεσ οβϖιουσ. Τηε 

ινφλυενχε χαλχυλατεδ ιν µοδελ ΜΙΙ ισ αλµοστ ινσιγνιφιχαντ ωηερεασ ιν µοδελ ΜΙ 

µαξιµυµ πυλλ ουτ στρεσσ ινχρεασεσ οϖερ 30 περχεντ χαυσεδ βψ τηε δεφορµατιον οφ 

τηε χονχρετε ελεµεντσ. Αδδιτιοναλλψ τηε βονδ στρεσσ ισ χαλχυλατεδ ασ 1.5 τιµεσ ασ 

ηιγη ασ φορ χασε (α) ατ ζ = 15µµ ωηιχη χαν βε εξπλαινεδ βψ τηε βουνδαρψ χονδι− 

τιονσ ανδ τηε ρεσυλτινγ στρεσσ φιελδ οφ χονχρετε. 

Bond stress, N/mm 2 

20 

18 

16 

14 

12 

10 



MI, case (b) 

MI, case (b), prestressed 

MII, case (b) 

MII, case (b), prestressed 

load 

8 

0 2 4 6 8 10 12 14 16 18 20 

embedment length z, mm 

Φιγυρε 9: Χοµπαρισον οφ χαλχυλατεδ βονδ στρεσσ ατ µαξιµυµ πυλλ ουτ λοαδ 

122


Influence of the reinforcement strain on the bond performance 

Τηε ρεσυλτσ οφ τηε χαλχυλατιονσ οφ τηε πυλλ ουτ στρεσσ ϖερσυσ σλιπ φορ χασε (χ) 

αρε σηοων ιν Φιγυρε 10. Φορ τηε µοδελ ΜΙ τηε µαξιµυµ λοαδ ισ ϕυστ αβουτ 15 

περχεντ ηιγηερ τηαν φορ χασε (α) ανδ λοωερ τηαν φορ χασε (β). Αλσο τηε βονδ στρεσσ 

οϖερ τηε εµβεδµεντ δεπτη ισ λοωερ φορ χασε (χ) χοµπαρεδ το χασε (α), ασ σηοων 

ιν Φιγυρε 12. 

load, N 

load, N 

1200 

1000 

800 

600 

400 

200 


MI, case (c) 

MII, case (c) 

0 

0,0 0,1 0,2 0,3 0,4 0,5 0,6 

slip, mm 

1200 

1000 

800 

600 

400 

200 

Φιγυρε 10: Χαλχυλατεδ πυλλ ουτ λοαδ ϖερσυσ σλιπ 


MI, case (c), prestressed 

MII, case (c), prestressed 

0 

0,0 0,1 0,2 0,3 0,4 0,5 0,6 

slip, mm 

Φιγυρε 11: Χαλχυλατεδ πυλλ ουτ λοαδ ϖερσυσ σλιπ οφ πρεστρεσσεδ σπεχιµεν 

123 



Ιν τηε χαλχυλατεδ βονδ στρεσσ σλιπ χυρϖεσ φορ πρεστρεσσεδ ρεινφορχεµεντσ, 

σηοων ιν Φιγυρε 11, τηε ινφλυενχε οφ τηε στεελ στραιν ον τηε βονδ περφορµανχε ισ 

αλσο οβϖιουσ. ∆υε το τηε πρεστρεσσινγ οφ αδηεσιϖε τψπε, ατ ινιτιαλ στατε τηε ρειν− 

φορχεµεντ λατεραλ στραινσ αρε νεγατιϖε (χοµπρεσσιον). Χονσεθυεντλψ, τηε χοντραχ− 

τιον οφ ρεινφορχεµεντ ισ λεσσ τηαν ιν τηε χασε οφ υνπρεστρεσσεδ ρεινφορχεµεντ ανδ 

τηερεφορε ιτ δοεσ νοτ ρεσυλτ ιν συχη α λαργε ρεδυχτιον οφ βονδ στρεσσεσ. Φιγυρε 12 

σηοωσ τηε ινφλυενχε οφ ∀Χ ον τηε βονδ στρεσσ φορ υνπρεστρεσσεδ ανδ πρεστρεσσεδ 

ρεινφορχεµεντ. 

Bond stress, N/mm 2 

20 

18 

16 

14 

12 

10 



MI, case (c) 

MI, case (c), prestressed 

MII, case (c) 

MII, case (c), prestressed 

load 

8 

0 2 4 6 8 10 12 14 16 18 20 

embedment length z, mm 

Φιγυρε 12: Χοµπαρισον οφ χαλχυλατεδ βονδ στρεσσ ατ µαξιµυµ πυλλ ουτ λοαδ 

Comparison of calculations and tests of textile reinforced elements in a 

bending test 

Ιν Φιγυρε 13 τηε ρεσυλτ οφ α φουρ−ποιντ βενδινγ τεστ ισ χοµπαρεδ ωιτη τηε ρε− 

συλτσ οφ τηε ΦΕ χαλχυλατιονσ. Τηε σπαν οφ τηε πλατε ωασ 250 µµ ανδ τηε λοαδ ωασ 

αππλιεδ ατ τηε τηιρδ ποιντσ. Τηε τεστ σπεχιµεν ωασ α εποξψ ιµπρεγνατεδ χαρβον 

τεξτιλε ρεινφορχεδ χονχρετε πλατε (300µµ ξ 60µµ ξ 10µµ) ωιτη α φινε γραιν 

χονχρετε (φχ ≈ 80ΜΠα). Τηε ινπυτ παραµετερσ φορ τηε χαλχυλατιον οφ τηε βενδινγ 

τεστ ωερε χαλιβρατεδ ατ δουβλε σιδεδ πυλλουτ τεστσ. Φορ δεταιλσ σεε /ΚΡ⇐ΓΕΡ, 

2002/. Φορ τηε χαλχυλατιον χονχρετε ισ µοδελλεδ βψ τηε µιχροπλανε µοδελ. Τηε 

αγρεεµεντ βετωεεν σιµυλατιον ανδ εξπεριµενταλ ρεσυλτσ ισ γοοδ. Ηοωεϖερ, χοµ− 

παρεδ το τηε τεστ ρεσυλτσ, τηε νυµεριχαλ ρεσυλτσ σηοω ιν τηε ποστ πεακ ρεγιον µορε 

δυχτιλε βεηαϖιουρ. 

124


Load, N 

1200 

1000 

800 

600 

400 

200 

Calculated, Case (a) 

Calculated, Case (c) 

test data 

0 

0 5 10 15 20 25 

Displacement, mm 

Φιγυρε 13: Λοαδ−δεφλεχτιον χυρϖε φορ α χαρβον ρεινφορχεδ ελεµεντ υνδερ µονοτονιχ λοαδινγ 

Ασ χαν βε σεεν φροµ Φιγυρε 14, τηε στραιν ιν ξ διρεχτιον (δαρκ ζονεσ ινδιχατε 

χραχκσ) σηοω α γοοδ αγρεεµεντ ωιτη τηε χραχκ διστριβυτιον οβσερϖεδ ιν τηε εξ− 

περιµεντ. Ιτ ηασ το βε νοτεδ τηατ τηε χραχκ διστριβυτιον ιν τηε τεστεδ σπεχιµεν ισ 

γρεατλψ ινφλυενχεδ βψ τηε τρανσϖερσε τεξτιλε ρεινφορχεµεντ, ωηιχη λεαδσ το λεσσ 

χραχκσ βυτ α ηιγηερ χραχκ ωιδτη. 

1 

Y 

Z 

X 

utput Set: MASA3 pbzfCEP6072 

eformed(20.26): Total nodal disp. 

χραχκσ 

τεξτιλε 

ρεινφορχεµεντ 

0.03 

0.0287 

0.0275 

0.0262 

0.025 

0.0238 

0.0225 

0.0212 

0.02 

0.0187 

0.0175 

0.0163 

0.015 

0.0138 

0.0125 

0.0113 

0.01 

0.00875 

0.0075 

0.00625 

Φιγυρε 14: Χοµπαρισον οφ πρινχιπλε στραινσ οφ χαλχυλατεδ µοδελ ιν ξ διρεχτιον ατ µαξιµυµ 

λοαδ ανδ χραχκ διστριβυτιον οφ τεστεδ σπεχιµεν. 

125 

0.005 



Ιν Φιγυρε 15 τηε στραινσ οφ τηε χονχρετε ελεµεντσ ιν ψ διρεχτιον (ηοριζονταλ 

χραχκσ) αρε σηοων ανδ χοµπαρεδ ωιτη σπεχιµεν. Τηε δαρκ ζονεσ ρεπρεσεντσ 

χραχκεδ χονχρετε. Ιν τηε εξπεριµεντ αλµοστ τηε σαµε χραχκ διστριβυτιον ανδ τηε 

σαµε φαιλυρε µοδε ωασ οβσερϖεδ. 

L1 

C1 

Y 

Z 

Output Set: MASA3 pbzfCEP6080 

X 

Deformed(22.23): Total nodal disp. 

Conto 

r A rg E stra 

Φιγυρε 15: Χοµπαρισον οφ τεστεδ σπεχιµεν αφτερ τεστ ανδ πρινχιπλε στραινσ οφ χονχρετε ελε− 

µεντσ ιν ψ διρεχτιον ατ µαξιµυµ λοαδ. 

CONCLUSIONS 

Α νεω δισχρετε βονδ µοδελ τηατ ισ βασεδ ον α βονδ στρεσσ−σλιπ ρελατιονσηιπ 

ηασ ρεχεντλψ βεεν ιµπλεµεντεδ ιντο α 3∆ φινιτε ελεµεντ χοδε. Τηε βονδ µοδελ 

αχχουντσ φορ τηε ινφλυενχε οφ ελαστιχ ανδ πλαστιχ ρεινφορχεµεντ στραινσ, τηε ινφλυ− 

ενχε οφ τηε ραδιαλ στρεσσ οφ τηε συρρουνδινγ χονχρετε ασ ωελλ ασ φορ τηε ινφλυενχε 

οφ τηε χψχλιχ λοαδ ηιστορψ ον τηε βονδ ρεσπονσε. 

Ασ σηοων ιν τηε νυµεριχαλ εξαµπλεσ, τηε τρανσϖερσε στρεσσ φιελδ ανδ τηε ρε− 

ινφορχεµεντ στραιν µαψ ηαϖε σιγνιφιχαντ ινφλυενχε ον τηε λοχαλ βονδ στρεσσ. Ιφ α 

ρεινφορχεµεντ ωιτη α ρουγη συρφαχε ισ υσεδ τηε λοχαλ βονδ στρεσσ ισ µαινλψ ινφλυ− 

ενχεδ βψ τηε ραδιαλ στρεσσ οφ τηε συρρουνδινγ χονχρετε. Ηοωεϖερ, τηε ινφλυενχε οφ 

τηε ρεινφορχεµεντ στραιν ινχρεασεσ ασ σµοοτηερ τηε ρεινφορχεµεντ συρφαχε ισ. 

Τηε παραµετερ ∀Χ ινφλυενχεσ τηε λοχαλ φαιλυρε οφ χονχρετε χλοσε το τηε βαρ 

νεαρβψ α χραχκ ωηερε ρελατιϖελψ ηιγη στρεσσεσ ιν ρεινφορχεµεντ αρε πρεσεντ. Τηε 

βονδ στρενγτη ισ ρεδυχεδ ανδ τηερεφορε χραχκ ωιδτη ανδ τηε διστριβυτιον οφ χραχκσ 

ισ αφφεχτεδ. Ιτ ισ ωελλ κνοων τηατ αλσο ψιελδινγ οφ στεελ ρεινφορχεµεντ ενλαργε τηισ 

εφφεχτ ωηιχη ισ αχχουντεδ φορ βψ ∀Σ ιν τηε βονδ µοδελ. 

126 

0.005 

0.00475 

0.0045 

0.00425 

0.004 

0.00375 

0.0035 

0.00325 

0.003 

0.00275 

0.0025 

0.00225 

0.002 

0.00175 

0.0015 

0.00125 

0.001 

0.00075 

0.0005 

0.00025 

0.


Τηε βενεφιτ οφ τηε πρεσεντεδ βονδ µοδελ βεχοµεσ οβϖιουσ ιφ ονε χονσιδερ 

διφφερεντ τψπεσ οφ ρεινφορχεµεντσ, ε.γ. διφφερεντ διαµετερ, συρφαχε στρυχτυρεσ ορ 

στεελ λυγσ ανδ στρεσσ στραιν προπερτιεσ. Ιτ ισ ασσυµεδ τηατ τηε γενεραλ παραµετερσ 

οφ τηε βονδ στρεσσ σλιπ−ρελατιον σηοων ιν Φιγυρε 1 µαινλψ δεπενδ ον τηε χονχρετε 

παραµετερσ ανδ χαν βε σετ το χονσταντ φορ α γρουπ οφ ρεινφορχεµεντ ελεµεντσ οφ 

τηε σαµε τψπε. Τηισ χαν βε φορ εξαµπλε α σετ οφ στεελ βαρσ οφ διφφερεντ διαµετερ 

ορ τεξτιλε ρεινφορχεµεντ τψπε ωιτη διφφερεντ Ψουνγ�σ µοδυλυσ βυτ αλµοστ τηε 

σαµε συρφαχε ρουγηνεσσ. 

Νεϖερτηελεσσ τηε δισχυσσεδ µοδελ ηαϖε το βε χαλιβρατεδ βασεδ ον α σεριεσ οφ 

διφφερεντ εξπεριµενταλ τεστσ ιν ορδερ το φινδ ουτ τηε ρεαλ ινφλυενχε οφ τρανσϖερσε 

στρεσσεσ ανδ στραινσ ον τηε βονδ ρεσπονσε ανδ τηυσ ον τηε στρυχτυραλ ρεσπονσε ασ 

ωελλ. 

REFERENCES 

Βαλαζσ, Γ.Λ. (1991): Φατιγυε οφ βονδ. ΑΧΙ Ματεριαλσ ϑουρναλ, 88 (6), 620−629, 

1991. 

Βραµεσηυβερ, Ω.; Βανηολζερ, Β.; Βρ⎫µµερ, Γ. (2000): Ανσατζ φ⎫ρ εινε ϖερειν− 

φαχητε Αυσωερτυνγ ϖον Φασερ−Αυσζιεηϖερσυχηεν. Βετον− υνδ Σταηλβετονβαυ, Νο. 

95, Ηεφτ 12, ππ. 702−706, 2000. 

ΧΕΒ Βυλλετιν 230 (1996): ΡΧ ελεµεντσ υνδερ χψχλιχ λοαδινγ. Στατε οφ τηε αρτ ρε− 

πορτ, Εδ. Βψ Τ. Τελφορδ, Τηοµασ Τελφορδ Σερϖιχε Λτδ, Λονδον, 1996. 

Χυρβαχη, Μ.; Ζαστραυ, Β. (1999): Τεξτιλβεωεηρτερ Βετον � Ασπεκτε αυσ Τηεοριε 

υνδ Πραξισ. Βαυστατικ−Βαυπραξισ 7, Μεσκουρισ (Εδ.), Βαλκεµα, Ροττερδαµ, 1999. 

Ελιγεηαυσεν, Ρ.; Ποποϖ, Ε.Π.; ανδ Βερτερο, ς.ς. (1983): Λοχαλ Βονδ Στρεσσ−Σλιπ 

Ρελατιονσηιπσ οφ ∆εφορµεδ Βαρσ υνδερ Γενεραλιζεδ Εξχιτατιονσ. Ρεπορτ 

ΥΧΒ/ΕΕΡΧ−83/23. Βερκελεψ: ΕΕΡΧ, Υνιϖερσιτψ οφ Χαλιφορνια, 1983. 

Κρ⎫γερ, Μ. (2001α): Πρεστρεσσεδ Τεξτιλε Ρεινφορχεδ Χεµεντ Χοµποσιτεσ. ΙΩΒ− 

Μιττειλυνγεν, ϑαηρεσβεριχητ 2000/2001, Υνιϖερσιτψ οφ Στυττγαρτ, Ινστιτυτε οφ Χον− 

στρυχτιον Ματεριαλσ (ΙΩΒ), 2001. 

Κρ⎫γερ, Μ.; Ρεινηαρδτ, Η.−Ω.; Φιχητλσχηερερ, Μ. (2001β): Βονδ βεηαϖιουρ οφ 

τεξτιλε ρεινφορχεµεντ ιν ρεινφορχεδ ανδ πρεστρεσσεδ χονχρετε. Ιν: Οττο−Γραφ− 

ϑουρναλ. ςολ. 12 2001, ππ. 33−50. 

Κρ⎫γερ, Μ.; Ρεινηαρδτ, Η.−Ω. (2001χ): Πρεστρεσσεδ τεξτιλε ρεινφορχεδ χεµεντ 

χοµποσιτεσ. Προχ. �11. Ιντερνατιοναλε Τεχητεξτιλ−Σψµποσιυµ φ⎫ρ τεχηνισχηε Τεξ− 

τιλιεν, ςλιεσστοφφε υνδ τεξτιλαρµιερτε Ωερκστοφφε�, Νο. 338, Φρανκφυρτ, Απρ. 2001. 

127 



Κρ⎫γερ, Μ.; Ξυ, Σ.; Ρεινηαρδτ, Η.−Ω.; Ο�βολτ, ϑ. (2002): Εξπεριµενταλ ανδ νυ− 

µεριχαλ στυδιεσ ον βονδ προπερτιεσ βετωεεν ηιγη περφορµανχε φινε γραιν χον− 

χρετε ανδ χαρβον τεξτιλε υσινγ πυλλ ουτ τεστσ. Ιν: Βειτργε αυσ δερ Βεφεστιγυνγσ− 

τεχηνικ υνδ δεµ Σταηλβετονβαυ (Φεστσχηριφτ ζυµ 60. Γεβυρτσταγ ϖον Προφ. ∆ρ.− 

Ινγ. Ρ. Ελιγεηαυσεν), ππ. 151−164, Στυττγαρτ, 2002. 

Λοωεσ, Λ. Ν.; Μοεηλε, ϑ.Π.; Γοϖινδϕεε, Σ. (2002): Α χονχρετε−στεελ βονδ µοδελ 

φορ υσε ιν φινιτε ελεµεντ µοδελλινγ οφ ρεινφορχεδ χονχρετε στρυχτυρεσ. Ιν πριντ, 

2002. 

Μαλϖαρ, Λ.ϑ. (1992): Βονδ ρεινφορχεµεντ υνδερ χοντρολλεδ χονφινεµεντ. ΑΧΙ 

Ματεριαλσ ϑουρναλ 89 (6), 711−721, 1992. 

Μενεγοττο, Μ.; Πιντο, Π. (1973): Μετηοδ οφ αναλψσισ οφ χψχλιχαλλψ λοαδεδ ρειν− 

φορχεδ χονχρετε πλανε φραµεσ ινχλυδινγ χηανγεσ ιν γεοµετρψ ανδ νονελαστιχ βε− 

ηαϖιουρ οφ ελεµεντσ υνδερ χοµβινεδ νορµαλ γεοµετρψ ανδ νονελαστιχ βεηαϖιουρ 

οφ ελεµεντσ υνδερ χοµβινεδ νορµαλ φορχε ανδ βενδινγ. Προχεεδινγσ οφ τηε 

ΙΑΒΣΕ Σψµποσιυµ ον τηε ρεσιστανχε ανδ υλτιµατε δεφορµαβιλιτψ οφ στρυχτυρεσ 

αχτεδ ον βψ ωελλ−δεφινεδ ρεπεατεδ λοαδσ, Λισβον, 1973. 

Ναµµυρ, Γ.; Νααµαν, Α. (1989): Βονδ Στρεσσ Μοδελ φορ Φιβερ Ρεινφορχεδ Χον− 

χρετε Βασεδ ον Βονδ Στρεσσ−Σλιπ Ρελατιονσηιπ. ΑΧΙ Ματεριαλσ ϑουρναλ, Νο. 86, 

ππ. 45−55, ϑαν./Φεβρ. 1989. 

Οηνο, Σ.; Ηανναντ, ∆.ϑ. (1994): Μοδελλινγ τηε στρεσσ−στραιν Ρεσπονσε οφ Χον− 

τινυουσ Φιβρε Ρεινφορχεδ Χεµεντ Χοµποσιτεσ. ΑΧΙ Ματεριαλσ ϑουρναλ, Νο. 91, 

ππ. 306−312, Μαρ. 1994. 

Ο�βολτ, ϑ.; Λι Ψ.; Κο�αρ, Ι. (2001): Μιχροπλανε µοδελ φορ χονχρετε ωιτη ρελαξεδ 

κινεµατιχ χονστραιντ. Ιντ. ϑ. οφ Σολιδσ ανδ Στρυχτυρεσ, 38, 2683−2711, 2001. 

Ο�βολτ, ϑ.; Λεττοω, Σ.; Κο�αρ, Ι. (2002): ∆ισχρετε βονδ ελεµεντ φορ 3∆ ΦΕ αναλψ− 

σισ οφ ρεινφορχεδ χονχρετε στρυχτυρεσ. Ιν: Βειτργε αυσ δερ Βεφεστιγυνγστεχηνικ 

υνδ δεµ Σταηλβετονβαυ (Φεστσχηριφτ ζυµ 60. Γεβυρτσταγ ϖον Προφ. ∆ρ.−Ινγ. Ρ. 

Ελιγεηαυσεν), ππ. 239−258, Στυττγαρτ, 2002. 

Ρεινηαρδτ, Η.−Ω.; Κρ⎫γερ, Μ. (2001): ςοργεσπανντε δ⎫ννε Πλαττεν αυσ Τεξτιλβε− 

τον. Προχ. �Τεξτιλβετον � 1. Φαχηκολλοθυιυµ δερ Σονδερφορσχηυνγσβερειχηε 528 

υνδ 532�, εδιτεδ βψ ϑ. Ηεγγερ, ππ. 165−174, Ααχηεν, 2001. 

Ψανκελεϖσκψ, ∆.Ζ.; Ρεινηαρδτ, Η.−Ω. (1987): Ρεσπονσε οφ πλαιν χονχρετε το χψ− 

χλιχ τενσιον. ΑΧΙ Ματεριαλσ ϑουρναλ 84 (5), 365−373 1987. 

128

Experimental realisation of a pretentious testing task on the field of pioneer bridge structures 

EXPERIMENTAL REALISATION OF A PRETENTIOUS TESTING 

TASK ON THE FIELD OF PIONEER BRIDGE STRUCTURES 

VERSUCHSTECHNISCHE REALISIERUNG EINER NICHT 

ALLTÄGLICHEN PRÜFAUFGABE AUS DEM BEREICH DER 

PIONIERBRÜCKENKONSTRUKTIONEN 

REALISATION D'UN ESSAI DE CHARGEMENT COMPLEXE D'UN 

PONTON DU GENIE MILITAIRE 

Wolfgang Harre 

SUMMARY 

An extraordinary test-setup and a pretentious test procedure is described for 

investigation of a ponton, submitted to the very manifold and complex loading 

conditions of pioneer bridge structures. 


Es wird der aufwendige Versuchsaufbau und die anspruchsvolle 

Versuchstechnik erläutert, um im Prüflabor die komplizierten 

Beanspruchungsverhältnisse mit allen Randbedingungen eines in eine belastete 

Schwimmbrücke eingebundenen Pontons nachzufahren, mit dem Ziel, die 

Reaktionen (Tragverhalten, Schwingfestigkeit) derartiger geschweißter 

Aluminium-Leichtbau-Konstruktionen zu untersuchen. 

RESUME 

Le dispositif et la procédure d'essai complexes servant à simuler en 

laboratoire les conditions de chargement très compliquées d'un ponton faisant 

partie d'un pont flottant sont décrites. 

KEYWORDS: Testing of Pioneer Bridge Structures, Aluminium-Bridge- 

Structures 

129 


W. HARRE 


The development of dismountable bridges (bridge systems, military 

bridges, pioneer bridges) requires, besides the extensive design work and 

detailed theoretical analysis, also experimental investigations on materials, 

structural details, complete substructures (modules) and even on complete 

bridge structures. Since many decades, the Otto-Graf-Institute is the leading 

testing institution on this field of dismountable bridges. 

A very important branch of the dismountable bridges are wet gap military 

bridging systems, the so-called floating bridges representing very efficient and 

universal useful structures to overcome big rivers, water surfaces and obstacles 

(caused for instance by catastrophes, floods etc.). 

In the following, it will be reported of a recently finished test project 

concerning floating bridge systems, which was outstanding pretentious if 

compared with the usual test projects, carried out commonly in the department 

2. 

2. TEST PROJECT 

Basic elements of floating bridges are generally the welded hollow box 

girders in aluminium, the so-called pontons or bays, which will – dependent on 

the respective demand – be coupled together in appropriate number to form in 

composite action a complete floating and load carrying road way, see the 

following pictures (Fig. 1 and 2). 

Fig. 1: Floating bridge in service 

130


Fig 2.: Cross-section of floating bridge (1 = Roadway ponton, 2 = Bow ponton) 

An excellent example of floating bridges, the so-called RIBBON-Bridge, 

was developed about 25 years ago in Germany by EWK (Eisenwerke 

Kaiserslautern). Meanwhile the RIBBON-Bridge is in service in 11 armies 

worldwide. Based on the positive experiences and the perfect performance all 

over the world in the past, even the US-Army was interested finally in this 

floating bridge. 

However in order to provide extensively the US-Army with the Ribbon 

Bridge, EWK had to satisfy some American proposals concerning a better 

handling and carrying capacity of the bridge system. At last, the Americans 

wished, that the successful demonstration of the demanded improvements 

should be realized by an appropriate full scale test in the Otto-Graf-Institute. 

So, in cooperation with EWK, a testing program was developed to prove 

the accomplishment of the requested improvements, quasi as a certification of 

the IMPROVED RIBBON BRIDGE (IRB). 

Essentially, the testing program included the static and dynamic loading of 

a complete ponton in a suitable special test set-up. Testing should be carried out 

in a procedure, which simulates all the load cases and load characteristics 

happening in practical use of the IRB. 

131 


W. HARRE 

To create that kind of most unfavourable and harmful loading situations in 

the ponton means concretely to apply bending moments and longitudinal as well 

as transverse loads in the same way and magnitude as it occurs, when a battle 

tank MLC 70 either stands on the floating bridge or passes the bridge, like 

demonstrated in the following diagram (Fig 3.). 

F9 

F10 

F1 

F2 F 

F3 F 

F4 F 

Ponton 

Fig. 3: Tank loading in reality and through laboratory simulation 

132 

F5 F 

F6 F 

F7 

F8


3. EXPERIMENTAL REALIZATION 

3.1 Test set-up 

There were mainly two problems to be solved: 

a) Installation and program-operation of a relatively high number 

(13) of hydraulic jacks with different capacities (50 kN to 2 MN) 

b) Application of high bending moments (partly > 2 MNm by 

means of very high horizontal forces) 

Even the comparatively abundant and well assorted equipment of the 

Department 2 of the Otto-Graf-Institute was not able and sufficient to solve 

satisfactorily the two problems in a direct way. Especially the wanted number 

and capacity of the hydraulic jacks (at least 4 jacks with more than 2 MN) 

represented a considerable challenge. So, some reflection was necessary to find 

a way for the experimental realization, using the available equipment. 

The central idea of the solution was the consistent application of the leveraction 

(Hebelgesetze): The mainly in pairs acting high horizontal forces with 

opposite signs could be replaced by a lever-structure as shown basically in the 

following sketch. 

F11 

F12 

F1- 6 

Ponton 

F13 

Fig. 4: Forces on the ponton 

133 


W. HARRE 

F7 

a 

jack > 1MN 

F8 

Ponton jack > 1MN Ponton 

b c 

task solution 

Fig. 5: Solution of the testing task 

jack < 1MN 

Proceeding that way, several profitable effects could be achieved: the 

number of the required jacks was reduced from 13 to 5, the capacity of the used 

jacks could be adapted perfectly to the jacks available in the department 2 by 

corresponding choice of the lever-ratios b:c and last not least – this is very 

important with regard to the test set-up – the introduction of the lever structures 

opens the possibility to anchor the initially horizontal forces now vertically 

either in the strong floor directly or by means of test frames at hand indirectly. 

The anchorage of high horizontal forces is – as experience shows – on principle 

very difficult and expensive in a laboratory, because these forces have to be 

turned round sooner or later to pass and anchor them finally into the floor. 

After the concept of the experimental realization was found, the proceeding 

was evident: after checking the available and suitable jacks in the department, 

the different lever-ratios were calculated for the different loading points. After 

that, all the other details were designed. Then all parts were manufactured by 

EWK together with the ponton to be tested. 

The following pictures will try to give an impression of the complicate and 

complex test set-up: 

134


Fig. 6. Testing arrangement on the strong floor 

135 


W. HARRE 

Fig. 7.a: Mid section of testing structure 

Fig. 7.b: End section of testing structure 

136


Fig. 7.c: Oil supply system 

Fig. 7.d: Complex multi-dimensional loading arrangement 

137 


W. HARRE 

3.2 Test run 

The loading program required the independent, however exactly balanced, 

synchronous controlling of altogether 5 jacks for static as well as for dynamic 

running. The following systematic presentation of the loading functions 

illustrates the dynamic test run (Fig. 8). 

Tension 

Tension 

Compression 

Compression 

Tension 

Compression 

Load-time-diagram 

Test run 

Fig. 8: Loading sequence 

138 

jack 1und 5 

jack 2 

jack 3 

jack 4 

time 

time 

time 

time


The implementation of this working load test procedure supposed the 

electronic coupling of the controlling units (S-59 Regler) of all the jacks. 

The exact run down of the whole test program was realized by means of a 

„managing“ computer, which directed the different controlling units according 

to the test program. In certain intervals, that is at times after reaching 

reconceived numbers of cycles, the test program also provided breaks in the 

dynamic loading. During these breaks, different static extreme load 

configurations were tested. All the measurements (loads, displacements, strains) 

happened automatically by a multipoint measuring system. The data were stored 

on CD for further evaluation. The described test set-up and equipment allowed a 

dynamic loading frequency of 0.4 Hz. The estimated lifetime of the bridge was 

50 000 cycles, so that the bare net time for carrying out the dynamic tests 

amounted to ca. 56 hours. 

The main result finally was, that the test specimen, will say the ponton 

(bay), as well as the test set-up itself passed the test procedure successfully. 

Apart from some insignificant cracks in the welds on uncritical structural points 

of the ponton, no serious damage could be observed on the test specimen. The 

test set-up also showed a perfect performance with regard to function and 

reliability. 

As a final statement, it can be concluded, that the experimental realization 

of this pretentious testing task on the field of pioneer bridge structures was an 

complete success for EWK and OGI. 

139 


W. HARRE 

140

Restoration of the sarcophagus of Duke Melchior von Hatzfeld 

RESTORATION OF THE SARCOPHAGUS OF DUKE MELCHIOR 

VON HATZFELD – THE ACCOMPANYING SCIENTIFIC AND 

TECHNICAL INVESTIGATIONS 

SCHADENSURSACHEN DES ZERFALLS DES HATZFELD- 

SARKOPHAGS UND ENTWICKLUNG EINER 

RESTAURIERUNGSMETHODE 

RESTAURATION DU SARCOPHAGE DU DUC MELCHIOR VON 

HATZFELD - INVESTIGATIONS SCIENTIFIQUES ET TECHNIQUES 

Gabriele Grassegger 

SUMMARY 

The article shows the investigations that led to the cause of decay of the 

alabaster sarcophagus and new methods for the restoration of the resin 

impregnated piece of art. The main reason was thermal decomposition of the 

gypsum and rapid rehydration accompanied by mismatching properties of the 

resin. The restoration used different formulas of cold-hardening PMMA resins 

combined with fillers and special coatings for 5 different steps of structural 

strengthening, adhesion, gluing of cracks, reshaping and retouching. The 

restoration has been successfully completed by a team of restorers. 


Der Alabaster-Sarkophag des Grafen von Hatzfeld (in Laudenbach) zeigte 

durch problematische Restaurierungen schwere Schäden, deren Ursachen 

festgestellt wurden. Die Hauptprobleme waren Wasserabspaltungen und Zerfall 

des Gipses, Spannungen durch Rückhydratisierung und andere physikalische 

Eigenschaften des Harzes, das zur Tränkung verwendet worden war. Es wurden 

dazu passende Restaurierungsmethoden entwickelt, die auf einer 5-stufigen 

Behandlung mit kalterhärtenden PMMA-Harzen beruhen. Die Behandlungen 

haben die Funktionen: strukturelle Festigung, Rißverklebung, Rißverfüllung und 

Antragungen, um die alte Form der Ornamente wieder herstellen zu können. Das 

Verfahren ist durch ein Restauratorenteam erfolgreich umgesetzt worden und 

der Sarkophag konnte wieder aufgestellt werden. 

141 


G. GRASSEGGER 

RESUME 

Suite à une restauration problématique, le sarcophage en albâtre du duc 

Melchior von Hatzfeld avait de sévères dégradations, dont les causes furent 

déterminées. Les causes principales étaient la décomposition thermique du 

plâtre, les tensions mécaniques dues à une réhydratation rapide, ainsi que les 

propriétés physiques mal adaptées de la résine utilisée. Nous avons développé 

une nouvelle méthode de restauration basée sur un traitement en cinq phases 

avec des résines PMMA durcissant à froid. Le traitement avait pour buts: le 

renforcement de la structure, le colmatage des fissures et le remodelage des 

ornements. La restauration a été accomplie avec succès par une équipe de 

restaurateurs, le sarcophage est à nouveau exposé. 

KEYWORDS: Restoration, alabaster, sarcophagus, conservation 


The sarcophagus of the late Duke Melchior von Hatzfeld was created in 

1659 by the famous stonemason Archilles Kern from Forchtenberg 

(Unterfranken). In the year 1657 Melchior von Hatzfeld had been the liberator 

of Krakau against the Swedish army sent by the German Emperor. The 

sarcophagus was finely carved out of a famous alabaster coming close by 

Forchtenberg. It shows the Duke in a suit of armour on the cover plate and 

scenes of his battles on the sides. Because of his legacy Archilles Kern created 

two tombs with very similar sarcophagus one in Prausnitz (Silesia) and one in 

Laudenbach (Hohenlohe) in a little chapel in the mountains, called Bergkirche 

(fig. 1). 

142


Figure 1: The Hatzfeld Sarcophagus after the successful restoration and reconstruction in the 

Bergkirche chapel in Laudenbach (picture by Georg Schmid, restorer). 

143 


G. GRASSEGGER 

Figure 2: Severe decay forms like warping, cracks and lamellar disintegration after the false 

restoration on one of the plates showing scenes from a battle 

(picture by Georg Schmid, restorer). 

144


Fig. 3: Stalk-like gypsum structure of very fine-grained alabaster formation impregnated and 

coated with PMMA. The large bubbles (vacuoles, below) are occasional places where air was 

trapped. Sample taken from the dog sculpture at a depth of approx. 10 cm (SEM picture). 

Fig. 4: Coarse gypsum crystal (left) with synthetic resin coatings. The ring-shaped objects are 

cavities in the crystal which are lined with a film of resin. The flaky body on the right is 

probably a newly developed anhydrite with a porous structure. Sample taken from a depth of 

approx. 10 cm (SEM picture). 

147 



Table 1: Composition of the restoration materials employed 

(Restoration expert Georg Schmid and team, Möglingen). 

Application Description/Recipe 

Adhesive mortar, fine formula for repair and 

shaping 

Injection mortar for adhesive and bridging of 

the cracks 

Structural strengthening of the damaged 

alabaster regions and pre-treatment of sides of 

cracks 

FM1= Acrylic resin suspension (Motema 

WPC) binder blended with Lenzin (natural 

gypsum) filler to a stiff, doughy consistency. 

I2= 1 part*) resin (Motema injection 

PMMA 220) + 1% by weight (in relation to 

the resin) peroxide catalyst plus 2 parts glass 

pellets

G. GRASSEGGER 

Table 2: Pressure resistance in various strengthened gypsum samples (block, dimensions 

approx. 5x5x2.5 cm, test following standard DIN EN 1926, test vertical to height). 

Sample Treatment of sample Bulk density 

[kg/dm³] 

Breaking load 

[N] 

Compressive strength 

[N/mm²] 

1 strengthened 2.26 23480.00 17.78 

2 strengthened 2.22 40750.00 30.57 

3 strengthened 2.20 34910.00 25.25 

4 strengthened 2.22 59850.00 46.61 

5 strengthened 2.21 25630.00 19.18 

6 strengthened 2.20 60050.00 44.96 

Average 2.22 40778.33 30.72 

A gypsum, freshly quarried 2.16 29870.00 19.41 

B gypsum, freshly quarried 2.24 28560.00 20.41 

C gypsum, freshly quarried 2.21 21120.00 17.88 

Average 2.21 26516.67 19.23 

Based on these findings, the result of strengthening could be rated very 

good. There was a substantial increase in resistance to pressure, from 19 to 30 

N/mm 2 on average, equivalent to a rise of c. 50%. An even greater improvement 

in strength was to be expected in the case of disintegrating gypsums like those in 

the sarcophagus, since a larger quantity of saturating material could be absorbed 

and residual strength had dropped to almost zero because of the destruction 

process. 

150


Figure 5: Before the restoration, small putto statue with most severe damage as warping, 

cracks and almost complete disintegration (picture by Georg Schmid). 

Figure 6: The same statue after restoration and treatment with 4 steps according to the 

methods proposed (picture by Georg Schmid). 

151 


G. GRASSEGGER 

Sealing the cracks in the alabaster 

Numerous cracks in the alabaster and the open joints between the sections 

had to be tied positively without visible changes. For this, original alabaster 

material was glued together with various mixtures based on Motema 220 (Table 

3). 

Table 3: Tensile strength of gluing of two pieces of gypsum with various adhesives based on 

Motema 220 (measured in accordance with the DIN EN 12 372 standard). 

Sample Breaking load, 

total (N) 

Tensile strength 

(N/mm 2 ) 

K1-1 gluing with filled resin*) 1531 0.61 

K1-2 gluing with filled resin 1542 0.61 

K1-3 gluing with filled resin 2044 0.82 

K1-4 gluing with filled resin (premature failure due 

to crack in gypsum) 

144 0.06 

K2 PMMA resin + 1% hardener, unfilled 2049 0.82 

K3 PMMA resin + 1% hardener, unfilled 822 0.33 

Average 1355 0.54 

*) comparable to Recipe I2 with the addition of 1% Aerosil (precipitated silicic acid) as a filler. 

The findings showed that all tensile tests on glued samples (both filled and 

unfilled adhesives) have a high level of tensile strength, higher than that of the 

stone itself. This is shown by the path of the fracture in the stone itself, i.e. a socalled 

cohesion fracture occurs in the stone. 

UV resistance and ageing tests on the finished mixtures 

For the sake of certainty, to check the durability of the finished mixtures 

they were tested for UV resistance. The plan was to expose all recipes that might 

be considered for use (20 in all) so as to rule out future changes. 

A climate simulator of the global UV testing type, model UV 200 RB/20 

DU, system Weiss, construction type BAM, was used. In this case, only UV 

radiation in long-term climatic conditions corresponding to the room climate 

was used. UV exposure took place in 2 cycles for a total of 300 hours. The 

samples were inserted vertically and half of each was covered with opaque foil 

(cf Figs. 4 and 5). 

UV radiation was by means of fluorescent lamps that approximate the 

short-wave part of sunlight. In particular, radiation simulates the high-energy 

152


UV-A and UV-B rays (λ = 300–420 nm) that could trigger photo-oxidation. The 

combination of fluorescent lamps employed corresponded to the spectral 

distribution as per Method B of DIN 53 384 E. 

Results of UV ageing 

No kind of UV ageing or other damage was found to result from storage in 

the room climate conditions and UV radiation. In this respect, the restoration 

materials must be described as durable and stable. 

CONCLUDING REMARKS 

These extensive tests created excellent conditions for the tomb’s lasting 

restoration. Skilful implementation by the team of restorers led by t Mr. Georg 

Schmid/Stuttgart Möglingen reinstated the tomb to its former beauty (see Fig. 1 

and 5). 

By way of additional protection, the grave chapel containing the 

sarcophagus is to be air-conditioned with the aim of avoiding alternating strains 

in future. For the reasons stated at the beginning, a constant climate of approx. 

10°C and a maximum of 50% relative humidity is to be aimed for. 

ACKNOWLEDGEMENT 

Thanks to the whole group of people who participated for the long period 

of investigations and trials until the sarcophagus could be restored, especially to 

Mr. Otto Wölbert and Mr. Meckes from the LDA who were the driving force of 

the project and never gave up. 

The whole history of the piece of art and it’s restoration will be published 

in the upcoming issue of the “Nachrichtenblatt der Denkmalpflege in Baden- 

Württemberg”, 4/2002 by a team of authors Otto Wölbert (restoration history), 

Georg Schmid (restoration), Judith Breuer (art history), Robert Vix 

(Architecture) and Gabriele Grassegger (technical investigations). 

REFERENCES/INTERNAL REPORTS (SELECTION) 

Grassegger, G. (1987): Hatzfeld-Grabmal, Bergkirche Laudenbach - 

Untersuchung zur Schadensursache an einem Alabaster-Sarkophag nach einer 

Kunstharz-Volltränkung, Nr. D3 140 008/GR (LDA internal report dated 

18.9.1987) 

153 


G. GRASSEGGER 

Grassegger, G. (2001): Restaurierung des Hatzfeld-Grabmals, Mechanische 

Untersuchung von Festigungen und Probeklebungen auf Alabastergips“, Nr. 

32-804073 (internal report of the Otto Graf Institute, Research and Testing 

Establishment for Building and Construction [FMPA] dated 2.7.2001) 

Grassegger, G. (2002): Restaurierung des Hatzfeld-Grabmals – Test der UV- 

Alterungsbeständigkeit bei den fertigen Restaurierungsmateralien (Kittmörtel, 

Klebungen, Injektagen und Retouchen), Berichtsnummer: 32 804 073 000-2 

(FMPA internal report dated 3.5.2002). 

154

Geotechnical Aspects and Observations of a Quarry Reclamation 

GEOTECHNICAL ASPECTS AND OBSERVATIONS OF A QUARRY 

RECLAMATION 

GEOTECHNISCHE ASPEKTE BEI DER WIEDERVERFÜLLUNG 

EINES STEINBRUCHS 

ASPECTS GEOTECHNIQUES DU REMPLISSAGE D'UNE CARRIERE 

Hermann Schad, Geoffrey Gay 

SUMMARY 

A disused quarry was refilled with mainly cohesive soil from excavations 

from the local area. During the refilling slip movements took place. The 

stabilisation methods used and the measurement and analysis of the movements 

that took place during the filling are described. 


Ein ausgebeuteter Steinbruch sollte mit bindigem Material aus der 

Umgebung – überwiegend Löss- und Verwitterungslehmen – verfüllt werden. 

Bei der Verfüllung traten trotz der Stabilisierungsmaßnahmen 

(Sandwichbauweise und Geokunststoffbewehrung) Rutschungen und größere 

Bewegungen auf. Eine Ergänzung dieser Maßnahmen durch Betonscheiben und 

einen Schotterfuß reduzierte die Verschiebungsgeschwindigkeit auf das bei 

Erddeponien übliche Maß. Durch die Langzeitbeobachtungen wurde es möglich, 

ein Kriechgesetz für die Bewegungen anzugeben. 

RESUME 

Une carrière désaffectée a été remplie avec des excavations de la région, 

principalement des sols cohérents. Pendant le remplissage, des glissements ont 

eu lieu. Les méthodes de stabilisation employées sont décrites, ainsi que les 

mesures et l'analyse des déplacements qui ont eu lieu pendant le remplissage. 

KEYWORDS: Limestone quarry, slip movement, stabilisation, refilling, 

geotextiles 

155 


H. SCHAD, G. GAY 


In 1985 it was decided to refill and recultivate part of a limestone quarry in 

the south west of Germany near Neuffen in the state of Baden-Württemberg. 

The quarry had been used for the production of crushed limestone mainly for the 

use in road construction. An aerial photograph of this stage of the refilling is 

shown in fig. 1. 

Fig. 1: Aerial photograph of first stage of refilling 24.4.1988 

In 1989 it was decided to refill a further part of the quarry. It was soon 

realised that the planned slope of 1:2 was not possible using conventional earth 

works construction methods so an arched retaining wall was planned at the foot 

of the slope and the fill was to be reinforced using geotextiles. The refilling was 

to be carried out using mainly cohesive soil from excavations in the vicinity. A 

cross-section of the planned refilling is shown in fig. 2. Later the slope was 

flattened to 1:2.5. 

156


Fig. 2: Cross-section of second stage of the refilling 1990 

2 SLIP MOVEMENTS AND STABILISATION 

In November 1997 it was noticed that relatively large slip movements must 

have taken place because a natural stone wall surrounding a manhole had been 

deformed considerably. (see fig. 3). The refilling was not complete at this time. 

Fig. 3: Deformed natural stone wall 

157 



It was therefore decided to set up a grid of measuring points to observe the 

movements of the slope. The grid points are shown in fig. 4. 

Fig. 4: Plan of grid points used between 24.11.97 and 21.11.00 

The first measurement took place in November 1997. After the first two 

measurements with a weeks difference between them it appeared that the 

deformation rate was slowing down. It had reduced from 9mm/d to 5mm/d. In 

the third week however the speed increased to 20mm/d so it was decided to 

increase the factor of safety by stabilising the foot of the slope using concrete 

buttresses with crushed rock between them as shown in figs. 5 and 6. 

158


Fig. 5: Section through concrete buttress 

Fig. 6: Plan of concrete buttresses with crushed rock filling 

159 



The concrete buttresses were used as a supporting element and the crushed 

rock as a drainage. Grid point measurements on the 23.12.97 showed a further 

increase in the deformation rate (37mm/d). It was therefore decided to fill the 

volume between the concrete buttresses with a well graded crushed rock instead 

of the coarse crushed rock as planned. The deformation rate slowed down 

considerably (see Section 3). At first it was not clear whether this was due to a 

frost period between 21.1.98 and 4.2.98. 

A slope stability calculation showed that for a factor of safety of 1.0 the 

shear parameters in the horizontal direction have to be ϕ = 10.21° and c = 0 

kN/m² and in the vertical direction ϕ = 20 ° and c = 0 kN/m² (see fig. 7). 

Fig. 7: Elements for slope stability calculations 

Calculations with the Kinematical Elements Methods (Gussmann et al. 

2002) for the stabilised state showed that the increase in stability factor due to 

the concrete buttresses and the lowering of the water table by 2m was relatively 

small (0.04). In the long term however an increase in the shear strength due to 

consolidation and “age hardening” is to be reckoned with. Under the phenomena 

“age hardening” is understood that when a cohesive soil is placed, especially in 

wet weather, there is a relative large amount of water between the soil 

aggregates and the soil is very soft or even “liquid”. In the course of time this 

160


The decisive deformation kinematics as derived from the deformation 

measurements between 24.11.97 and 8.2.98 are shown in fig. 4. It can be seen 

that the part which is reinforced with geotextiles moved horizontally as a block. 

The displacement rates of the points 16 and 17 (fig. 4) are characteristic for 

the lower part of the slope. The average deformations and the resulting 

deformation rates are shown in the following diagram. 

Fig. 9: Displacements of the lower part during the first 72 days 

After construction was completed a horizontal deformation of 16mm in the 

course of 1.5 years was measured at grid point 5 (see fig. 8). The vertical 

deformations of this point during the same time interval were 36mm. The 

maximum settlement measured on the “plateau” was 48mm. At the foot of the 

slope the maximum deformations at grid point 10 were horizontal 14mm and 

vertical 3mm. Of special note is the similarity of the settlements of the grid 

points 1 to 5 on the “plateau”. Inside 1.5 years (21.11.00 to 4.4.02) they were 

40mm, 48mm, 39mm, 36mm and 35mm as shown in fig. 10. 

162


Fig. 10: Displacements of the plateau in the second phase 

4 INTERPRETATION OF THE MEASUREMENTS 

In the following diagram (fig. 11) the average time deformation curves 

with the time on a logarithmic scale are plotted. It can be seen that up to 68 days 

after the start of the measurements slipping took place. After that the creep 

phase started. 

Fig. 11: Displacements as function of time 

163 



In the semi-logarithmic plot the creep phases approximate to straight lines 

which can be represented to a good approximation by the black lines with circles 

as points. Using logarithms to the base 10 the following relationships are 

obtained. This logarithmic creep is often observed in soil but a lot of other 

rheological models could be used (e. g. Schad/Breinlinger 1991). 

In the next 10 years horizontal deformations of 5 to 10 cm. are to be 

expected and that similar deformations are to be expected in the next 90 years. 

5 LITERATURE 

Gussmann, P.; Schad, H.; Smith, I. (2002): Numerical Methods in Geotechnical 

Engineering Handbook, Vol. 1, Ernst & Sohn Berlin, 437 - 479 

Schad, H.; Breinlinger, F. (1991): Numerical analysis of visco-elastoplastic soil 

behaviour considering large deformations. Proc. 10th European Conference 

on Soil Mechanics and Foundation Engineering, Florence/Italy, 255 - 260. 

164

Non-destructive detection of longitudinal cracks in glulam beams 

NON-DESTRUCTIVE DETECTION OF LONGITUDINAL CRACKS IN 

GLULAM BEAMS 

ZERSTÖRUNGSFREIE MESSUNG VON LÄNGSRISSEN IN 

BRETTSCHICHTHOLZ-TRÄGERN 

DÉTECTION NON DESTRUCTIVE DES FISSURES 

LONGITUDINALES DANS LES POUTRES DE BOIS LAMELLÉ 

COLLÉ 

Simon Aicher, Gerhard Dill-Langer, Thomas Ringger 

SUMMARY 

The paper reports on the detection and length characterisation of 

longitudinal cracks in glued laminated timber (glulam) beams by means of 

ultrasound (US) pulse transmission method. In the preliminary study one large 

glulam beam with a crack starting at one end-grain face and ending at about one 

third of total beam length has been evaluated. For the transmission 

measurements US pulses have been applied to the narrow faces of the beam, 

thus propagating parallel to cross-sectional depth perpendicular to fibre. The 

beam has been scanned by a transmitter / receiver pair of US transducers shifted 

along the longitudinal beam axis. The recorded full US wave signals were 

evaluated for three different scalar parameters being “time of flight”, “peak-topeak 

amplitude” and “first amplitude”. The comparison of the visual inspection 

with the US parameters, showing significantly different scatter ranges, yielded a 

satisfactory agreement with respect to the determination of crack length. The 

NDT crack detection based on the parameter “time of flight” was also 

satisfactory when the crack extended only over a part of the beam width, i. e. not 

being visually detectable from one of both side faces. The latter can be very 

important for in-situ inspection of beams in buildings with assumed or partial 

cracks. 

165 


S. AICHER, G. DILL-LANGER, T. RINGGER 


Der Aufsatz berichtet über den Nachweis und die Längenmessung von 

longitudinalen Rissen in Brettschichtholz (BSH) -Trägern mittels der Methode 

der Durchschallung mit Ultraschallpulsen. In der orientierenden Studie wurde 

ein großer Brettschichtholzträger mit einem Riss untersucht, wobei der Riss von 

einer Hirnholzfläche ausging und innerhalb des ersten Drittels der Trägerlänge 

endete. Für die Transmissionsmessungen wurden Ultraschallpulse auf die 

Schmalseiten der Träger aufgebracht, so dass diese sich parallel zur 

Querschnittshöhe und damit rechtwinklig zur Faserrichtung ausbreiteten. Der 

Träger wurde mit einem Ultraschall Geber / Empfänger-Paar in Richtung der 

Trägerachse abgerastert. Die aufgezeichneten vollständigen Ultraschallsignale 

wurden hinsichtlich dreier verschiedener skalarer Parameter ausgewertet: 

Signal-Laufzeit, Signalstärke und erste Amplitude. Der Vergleich der visuellen 

Charakterisierung mit den Ultraschallparametern, die jeweils deutlich 

unterschiedliche Streubreiten aufwiesen, ergab eine zufriedenstellende 

Übereinstimmung hinsichtlich der Bestimmung der Risslänge. Die 

zerstörungsfreie Risserkennung auf Grundlage des Parameters "Signal-Laufzeit" 

war auch dort noch zufriedenstellend, wo sich der Riss nur über einen Teil der 

Querschnittsbreite erstreckte, also auf einer der beiden Seitenflächen schon nicht 

mehr sichtbar war. Letzteres kann für die in-situ Beurteilung von Trägern in 

realen Bauwerken mit vermuteten oder teilweise vorhandenen Rissen sehr 

wichtig sein. 

RÉSUMÉ 

L’article traite de la détection et de la caractérisation des fissures 

longitudinales dans les poutres en bois lamellé collé, au moyen d’une méthode 

de transmission des impulsions ultrasons. Lors de travaux préliminaires, une 

poutre présentant une fissure commençant à l’une des extrémités et s’étendant 

jusqu’au tiers de la longueur totale a été étudiée. Les impulsions d’ultrasons sont 

appliquées sur les deux faces les plus étroites de la poutre et se propagent 

parallèlement à sa section, c’est à dire perpendiculairement aux fibres du bois. 

Les capteurs ultrasons (émetteur et transmetteur) balaient alors la poutre suivant 

son axe longitudinal. Le signal enregistré fournit trois paramètres différents: le 

temps de propagation du signal, son amplitude pic à pic et sa première 

amplitude. La comparaison entre les indications obtenues par la méthode des 

166


ultrasons et l’évaluation visuelle sont en accord quant à la détermination de la 

longueur de la fissure. La détection basée sur le seul paramètre « durée de 

propagation » est également satisfaisante lorsque la fissure ne s’étend que sur 

une partie de l’épaisseur, c’est à dire lorsqu’elle ne traverse pas la poutre de part 

en part. Ce dernier cas est très intéressant pour l’inspection in-situ des 

constructions présentant des fissures ou des risques de fissure. 

KEYWORDS: non-destructive testing, ultrasound, pulse transmission, crack in 

glulam beams, scalar ultrasound parameters 


Non-destructive evaluation of the state of integrity resp. of defects or 

partial damages in structural building elements generally represents an important 

issue. The capability of NDT based reliable assessment of components enhances 

the acceptance of building systems or materials, may affect safety factors and 

enables assessment of upgrading or rehabilitation works. Timber and glulam 

beams despite all positive aspects are prone to longitudinal cracks generally 

resulting from interaction of poor constructive detailing and unaccounted 

climatic stresses. Longitudinal cracks primarily occur at i) support areas due to 

interaction of shear stresses and climate and ii) at notches, holes and in apex 

areas of curved / tapered beams due to tension stresses perpendicular to grain 

bound to undue load actions and / or often climate stresses. Finally, 

iii) longitudinal cracks can occur from poor glue lines generally bound to 

trespass of open / closed curing times of the adhesives. 

The reliable assessment of the extent of damage and of the result of the 

repair works represent the two equally important aspects of the NDT assessment 

of damaged or upgraded construction elements. In many cases a visual 

inspection of the beams is very costly or almost impossible. Today reliable NDT 

tools for employment in the sketched area are missing for lumber / glulam 

beams being contrary to constructions with some other important building 

materials. The reason for this NDT lag consists i.a. in the anisotropy, 

inhomogeneity and the high damping characteristics of the natural material 

wood. 

167 



Timber Department of Otto-Graf-Institute being deeply involved in the 

assessment / expertises of damages, repair proposals and technical approval of 

rehabilitation works has started to focus on active NDT methods about a year 

ago. This paper gives some preliminary results of one of the on-going projects. 

It is reported on the detection and length characterisation of longitudinal cracks 

in glued laminated timber (glulam) by means of ultrasound (US) pulse 

transmission method. The US method has been chosen due to its sensitivity to 

impedance changes at discrete boundaries. Former literature known attempts in 

this field [KLINGSCH, 1991; KIMURA ET AL 1991] dealt with artificial defects, 

being rather thick saw cuts of defined length parallel to beam axis and over full 

cross-sectional width. The results and the practical relevance of the exclusive 

focus on such slots / cracks has been discussed controversially in the involved 

engineering community. In the investigation reported here a fully practice 

relevant crack resp. cracked beam was investigated. 

2. EXPERIMENTAL SET-UP 

The investigated specimen represented a part cut from a large beam with a 

round hole loaded until failure in bending, compare Fig. 1a. The beam had failed 

with two large cracks initiated at the hole periphery by high local stresses 

perpendicular to grain. The crack propagation was then driven by both, shear 

and tension stresses perpendicular to grain. 

The investigated NDT specimen (Fig. 1b) showed an open crack over full 

width b at end grain face Y closer to the former hole location and no visible 

crack at the opposite end grain face Z. Thus, despite disputable accuracy of 

visual inspection the crack must end within the specimen as the latter consisted 

of one massive piece. 

168

a) 

b) 

cracks at 

ultimate load 

originally 

tes ted beam 

h = 440 

end grain 

face Y 

33 

33 

100 

Y 

b = 120 

p/2 p/2 

L c 

y 

x 

T 

B 


l = 2100 

investigated 

NDT specimen 

Z 

l = 2100 

h = 440 

120 

Top narrow face (T) 

visible part of the crack 

Fig. 1a,b: Original location and dimensions of the employed NDT specimen. 

a) larger cracked beam from which the NDT specimen was cut 

out after failure of the beam 

b) view and dimensions of partially cracked NDT specimen 

169 

end grain 

face Z 

900 

"right" wide face II 

Bottom narrow face (B) 



The dimensions of the block were (width b × depth h × length l): 120 mm × 

440 mm × 2100 mm. Starting at end-grain face Y and proceeding in the 

longitudinal (x-) beam direction, the crack is visible at both side-faces I and II 

for the major part of the crack length (compare chap. 4). 

The ultrasonic NDT evaluation / assessment of the crack (length) was 

throughout performed by means of a pair of piezoelectric (US) transducers. 

Transmitter and receiver were positioned oppositely at mid-width of the narrow 

faces T and B of the beam specimen and aligned parallel to depth. Starting at the 

end-grain face Y with the opened crack the transducer pair was moved along 

beam length l with increments of �x = 50 mm towards the end grain face Z. At 

each position an ultrasonic pulse synthesized by a generator unit was put to the 

specimen by the piezoelectric transmitter. Figure 2 shows a schematic 

representation of the experimental set-up. The fixation of the transmitter and of 

the receiver differed. The transmitter was throughout fixed to the surface by a 

hot melt adhesive also serving as coupling agent. Contrary, the receiver was not 

glued to but applied to the surface by hand pressure without using any kind of 

coupling agent. 

At each location x a number of 25 repetitive measurements were performed 

in order to enable noise reduction. As the crack was not centred in the middle of 

the cross-sectional depth but much closer (~70 mm) to narrow face B it was 

questioned whether there might be an influence if the transmitter is at a closer or 

more remote distance to the crack. Therefore two test series S1 and S2 were 

performed with the transmitter first being at narrow side T and then at narrow 

side B. 

170

a) 

b) 

end grain 

face Y 

crack 

x 

narrow 

face T 

crack at end 

grain face A 

narrow 

face B 


l = 2100 

US-pulse 

generator unit 

ultrasound 

transmitter 

signal [V] 

3 

0 

-3 

narrow face T 

narrow face B 

0.0 time [ms] 0.5 

ultrasound 

receiver 

amplifier 

Fig. 2a,b: Schematic representation of the experimental set-up 

171 

end grain 

face Z 



3. NDT EQUIPMENT 

The generator unit (USG 20, Geotron Electronics), originally optimised for 

NDT of concrete, produced high voltage pulses with main frequencies between 

20 kHz and 350 kHz. The duration of a single pulse was less than 1 ms. 

The ultrasonic transducers (UPG-D 3037, UPE-D 3038, Geotron 

Electronics) used in the experiments were piezoelectric converters with a 

coupling-surface of 3 mm in diameter. The transducers showed a multi-resonant 

frequency characteristic with main peak values between 20 and 100 kHz. 

The received ultrasonic pulses have been amplified by a broadband 

amplifier (AM 502, Tektronix) with an amplification factor of 100 dB. The 

complete signals were recorded by a PC based transient recorder with 12 bit 

amplitude resolution and 20 MHz time resolution. 

4. VISUAL CHARACTERIZATION OF THE CRACK DIMENSIONS 

For correlation of the ultrasound NDT parameters with the length of the 

crack in longitudinal beam direction and with the crack opening, the dimensions 

of the crack were determined by visual inspection at both wide side faces I and 

II of the specimen. The crack openings were measured with a feeler gauge. 

Figures 3 a and 3 c give a schematic illustration of shape, position and 

dimensions of the crack according to the visual characterization and feeler gauge 

measurements at the two wide faces I and II, while Fig. 3 b shows a top view 

indicating the projected crack area. According to the visual findings the crack 

can be divided into three different sections. 

In section A (0 ≤ x ≤ 53 cm), the crack is characterized by measurable 

openings of 1.2 mm (x = 0) to 0.4 mm (x = 53 cm) at side face I; at side face II 

the respective dimensions are: 0.4 mm (x = 0) to 0.05 mm (x = 53 cm). 

In section B (53 cm < x ≤ 67 cm) a crack-opening was only measurable at side 

face I with openings in the range of 0.35 mm to 0.25 mm. At side face II, the 

closed crack was visible as a small displacement edge within the surface. 

Finally in section C (67 cm < x < 88.4 cm), the crack was still measurable 

at side face I with openings from 0.25 mm to 0.05 mm. The end of the crack at 

x = LC =88.4 cm almost coincides for measurable (0.05 mm) crack opening and 

visual inspection. At side face II, the crack is not visible at all. 

172


The true extension and shape of the crack front might well be somewhat 

ahead of x = LC what will be determined at the end of the ongoing experiments. 

5. CHARACTERIZATION OF SIGNAL-PARAMETERS 

Once an ultrasonic pulse is generated and applied to the narrow face of the 

glulam beam, it is proceeding through the specimen perpendicular to the 

direction of the glued lamellas, i.e. perpendicular to fibre direction and is 

detected by the receiver at the opposite surface. 

The recorded full wave signals purged from noise by multiple pulse 

measuring method were so far evaluated for three different scalar parameters, 

being: 

• “Peak-to-peak amplitude” (pp amplitude) of the signal, which represents 

the difference between the recorded absolute maximum and minimum of 

the complete signal. The parameter is correlated to the transmitted energy 

of the pulse. Figure 4 shows one exemplary wave signal, including the 

determination of the pp-amplitude. 

• “Time of flight” (TOF) of the signal is defined as the time lag between the 

external trigger edge given by the pulse generator and the on-set, i.e. the 

begin of the recorded signal. The signal-parameter TOF and also the 

below specified parameter |1 st 

a| are exemplarily depicted in Fig. 5 for the 

signal given in Fig. 4 now presented with a close-up at the begin of the 

signal. 

• “First amplitude” (1st a) of the signal is defined as the maximum (or 

minimum) amplitude of the first observable half cycle. In detail, for signal 

characterization, the absolute value of the first amplitude has been used. 

173 



a) 

b) 

c) 

h [cm] 

40 

20 

0 

40 

face Y 

20 

0 

L C 

A B C 

left face I 

0 70 140 x [cm] 210 

0 70 140 x [cm] 210 

x 

h [cm] 

40 

20 

0 

A B C “left“ face I 

53.0 cm 

L C 

67.0 cm 

A B C 

88.4 cm 

exact shape of crack front unknown 

“right“ face II 

end of measurable crack 

(openings > 0.05 mm) 

end of visible crack 

right face II 

0 70 140 x [cm] 210 

face Z 

beam depth 

h = 440 mm 

beam depth 

h = 440 mm 

beam width 

b = 120 mm 

Fig. 3a-c: Schematic illustration of the appearance of the crack at different faces of the 

specimen. The graphs 3a) and 3c) give measured crack lengths and crack 

openings (100-times enlarged) at the left and right wide side faces (I and II). Fig 

3b shows a projection of the crack area revealing the three crack sections A-C. 

174

signal [V] 

3.0 

1.5 

0.0 

-1.5 

close up 

in Fig. 5 


-3.0 

0.00 0.20 0.40 0.60 0.80 1.00 

pp amplitude 

time [ms] 

Fig. 4: Recorded signal with evaluation / definition of the “peak-to-peak amplitude” 

(pp amplitude) 

signal [V] 

0.4 

0.2 

0.0 

-0.2 

1 st a 

TOF 

-0.4 

0.00 0.05 0.10 0.15 0.20 0.25 

time [ms] 

Fig. 5 Evaluation / definition of “time of flight” (TOF) and of “first amplitude” (1st a); the 

graph represents a close-up of the recorded ultrasound pulse given in Fig. 4 

175 



6. RESULTS OF THE ULTRASOUND MEASUREMENTS 

The reproducibility of the signal parameters for repeated independent 

measurements at a specific location x (uncoupling and new coupling of the 

transducers for each measurement) differed considerably between parameter 

TOF on the one side and parameters pp and |1 st a| on the other side. In case of 

TOF in average an extremely small coefficient of variation (C.O.V.) of 0.3 % 

was obtained whereas for the parameters pp and |1 st a| the considerably higher 

C.O.V.’s were 21% and 24 %, respectively. For the reproducibility test a number 

of 10 repeated measurements have been evaluated. 

The major results of the performed preliminary investigations are compiled 

in Figs. 6 to 8, showing the signal parameters TOF, pp and |1 st a| along specimen 

axis x. In all graphs the results of the two test runs S1 and S2 with the alternative 

transmitter positions at narrow specimen sides T or B are given, and the mean 

value of both test runs is shown additionally. Further, the quality of the signal 

parameter reproducibility is indicated in all Figures by an error bar with a height 

of 2 times of the respective standard deviation (the error bars are not visible in 

Fig. 6 due to the very small C.O.V.’s). The visually determined crack length 

segments A, B and C are indicated in the graphs, too. Following the results are 

discussed in more detail. 

Figure 6 specifying “time of flight (TOF)” vs. beam axis x reveals almost 

throughout a steep decrease of parameter TOF along crack length segments A, B 

and C. It should be emphasized, that the TOF decrease in the investigated case is 

apparently not affected by the fact that the crack is not visually detectable at 

surface II in crack zone C. For positions x > LC a rather constant TOF value of 

253.2 ms is obtained. This gives a mean phase velocity in transverse direction to 

fibre of v90 = 0.44 / (253.2*10 -6 ) = 1738 m/s which is in good agreement with 

literature data [Buchur 1989] on phase velocities perpendicular to fibre of wood 

/ glulam made of European spruce. A comparison of test series S1 and S2 

indicates apart from one exception in the crack range A, that the TOF results are 

obviously not influenced by transmitter location closer resp. more far from the 

crack plane. 

176

TOF [µs] 

360 

330 

300 

270 

253.2 

240 

mean value 

of S1 and S2 


S1: transmitter at face T 

S2: transmitter at face B 

visually determined crack length, wide side face I 

visual crack length, wide side face II 

L C 

A B C 

0 30 60 90 120 150 180 210 

specimen axis x [cm] 

TOF mean value of 

uncracked specimen 

Fig. 6 Results of time of flight (TOF) measurements along the beam axis x. The x-axis gives 

the distance x [cm] of the transducers position to the end grain face Y. The crack was 

supposed, according to visual inspection, to end at x = 88.4 cm. 

Figure 7 shows the “peak-to-peak amplitude” (pp amplitude) of the 

transmitted signals. Qualitatively a slow increase of pp-amplitude values from 

mid- length of segment A through to C and the increase continues to about 20 

cm beyond C; into the visually uncracked part of the beam. Quantitatively high 

scatter of the measured data, especially in the uncracked section is observed. 

Exemplarily at a distance x = 160 cm from end grain face Y, certainly well 

ahead of the crack front, the pp-amplitudes exhibit a local minimum with values 

comparable to those measured within the crack at x = 70 cm (in section C). The 

transition from the cracked to the undamaged section is rather smooth without a 

clearly marked step in the pp amplitude course. 

177 



pp amplitude [V] 

15.0 

12.0 

9.0 

6.0 

3.0 

0.0 

vis. crack length 

wide side face I 

L C 

A B C 


wide side face II 



0 30 60 90 120 150 180 210 


mean value 

of S1 and S2 

Fig. 7 Results of peak-to-peak amplitude (pp amplitude) measurements along beam axis x. 

The x-axis gives the distance x [cm] of the transducers position to the end grain face 

Y. The crack was supposed according to visual inspection to end at x = 88.4 cm. 

Thus, the signal-parameter pp-amplitude does not allow a clear 

identification of the crack length. The relatively high uncertainty due to coupling 

conditions makes it even more difficult to quantitatively estimate the location of 

the crack tip. However, in spite of the scatter, the clearly visible trend of 

decreased attenuation for decreasing crack openings is not affected qualitatively. 

The presented results for the behaviour of the peak-to-peak amplitude in 

case of cracks can be compared to the observations of [KLINGSCH, 1991], where 

no damping of pp amplitudes has been measured in the case of saw-cuts. 

In Fig. 8 the results of |1 st a| along beam axis x are shown for the two 

performed test series S1 and S2 together with the boundaries of the different 

crack sections. 

178

| 1 st a | [V] 

0.80 

0.60 

0.40 

0.20 

0.00 


wide side face I 

L C 

A B C 



wide side face II 



0 30 60 90 120 150 180 210 


mean value 

of S1 and S2 

Fig. 8 Results of measurements of absolute values of the first amplitude (|1st a|) along the 

beam axis x. The x-axis gives the distance x [cm] of the transducers position to the end 

grain face Y. The crack was supposed according. to visual inspection to end at 

x = 88.4 cm. 

The measured course of the |1 st a| values can roughly be described as a step 

function with quite constant low values within all three sections A to C of the 

crack and a sharp increase at the assumed crack tip. It is interesting to note that 

in section C still strong damping of |1 st a| is observed, while the crack is solely 

visible at one side face of the specimen. 

Although the scatter among |1 st a| values within the undamaged part of the 

beam is significant, the results between cracked and uncracked parts of the beam 

are clearly separated, which is especially true for the mean values of the two test 

series with interchanged transmitter / receiver conditions. 

179 




The performed preliminary study on the feasibility of crack detection in 

glulam beams by means of ultrasonic pulse transmission method revealed 

promising results. 

All three evaluated scalar signal parameters, being time-of-flight (TOF), 

peak-to-peak-amplitude (pp-amplitude) and first amplitude (⏐1 st a⏐) showed 

significant correlations with the occurrence of the completely or partly visually 

detectable crack. 

Both, pp-amplitude and ⏐1 st a⏐ exhibited relatively high scatter among the 

values within the undamaged part of the beam accompanied by a quite poor 

reproducibility bound to the performed non-ideal coupling conditions. The 

pp-amplitude values showed a smooth transition zone from cracked to the 

uncracked sections of the beam. Contrary hereto the ⏐1 st a⏐ values exhibited a 

pronounced step indicating the end of the crack clearly. 

The TOF-results showed best reproducibility and a clear, although smooth 

change at the end of the crack. The different crack sections with one-sided 

throughout measurable crack openings respectively one sided first measurable 

and then visible crack opening were best represented by the course of TOF 

results. 

For all three characteristic signal parameters, no significant differences due 

to interchanged positions of transmitter and receiver were observed. Thus, no 

detection of the location of the crack with respect to depth direction could be 

performed, being in good accordance with the results of [KLINGSCH, 1989, 

1991]. 

Although feasibility of the applied NDT methods and evaluation for crack 

detection could be shown by the presented study, the results from only one 

exemplary specimen may not be generalized. Additional tests also with 

investigate different beam / crack configurations have to be performed to obtain 

a statistically more reliable data basis. 

In order to improve the presented ultrasonic method for applications in real 

structures, the coupling problem has to be solved and the feasibility for beams 

with realistic heights of about 1 to 1,5 m has to be shown. Advanced signal 

processing techniques for the evaluation in the frequency domain (i.e. Fourier- 

180


and Wavelet transforms) should be used for noise reduction, enhancement of 

resolution and defect sensitivity. 


The authors are very much indebted to Dr. Catherine Lidin (Collano AG, 

Switzerland) for the utmost valuable favour translating the abstract and title of 

this paper into technically and linguistically correct French. 

The financial support of German Science Community (DFG) via grant to 

Sonderforschungsbereich 381 "Characterisation of damage evolution in 

composite materials using non-destructive test methods" and hereby to subproject 

A8 "Damage and NDT of the natural fibre composite material wood" is 

gratefully acknowledged. 

REFERENCES 

KLINGSCH, W.: Zerstörungsfreie Lokalisierung äußerlich nicht sichtbarer 

Holzschädigung. Bauen mit Holz 6, 1989, pp. 421-423 

KLINGSCH, W.: Erarbeitung anwendungstechnischer Grundlagen zur 

zerstörungsfreien Qualitätsüberwachung von Holzleimbauteilen mittels 

Ultraschall. Forschungsbericht, 1991 

BUCUR, V.: Acoustics of wood. Boca Raton, New York, London, Tokyo, 1995, 

p. 121 

KIMURA, M., KUSUNOKI, T., OHTA, M., HATANAKA, K., KOZUKA, H., ITO, H.: 

Ultrasonic pulse test on glulam glued connection. Proc. Int. Timber Eng. 

Conf., part 2, London, 1991, pp. 2.250 – 2.257 

181 



182

Determination of local and global modulus of elasticity in wooden boards 

DETERMINATION OF LOCAL AND GLOBAL MODULUS OF ELAS- 

TICITY IN WOODEN BOARDS 

BESTIMMUNG DES LOKALEN UND GLOBALEN ELASTIZITÄTS- 

MODUL IN HOLZBRETTERN 

DETERMINATION DU MODULE D’ELASTICITE LOCAL ET GLO- 

BAL SUR DES PANNEAUX A BASE DE BOIS 

Simon Aicher, Lilian Höfflin, Wolfgang Behrens 

SUMMARY 

The paper reports on an efficient method for determination of the local 

modulus of elasticity by means of elongation/strain measurements. Further, the 

effect of local weak sections on the global modulus of elasticity determined by 

deflection measurement is revealed. The global modulus of elasticity computed 

on the basis of the partly extremely varying locally measured moduli of elasticity 

complies well with the globally measured MOE. 

The experimental investigations were performed with edgewise bent beech 

boards. First, the elongation /strain measurement method was verified exemplary 

with a board which was inflicted successively with artificial defects 

(holes). For each defect state the local and global moduli of elasticity were 

measured and the differences are discussed. Second, the variation of local 

modulus of elasticity and its high spatial correlation with the location of bending 

failure is shown exemplarily by means of four beech boards of a larger test series. 


Es wird über eine effiziente Methode zur Bestimmung des lokalen Elastizitätsmoduls 

mittels Längsverschiebungs-/Dehnungsmessungen berichtet. Desweiteren 

wird die Auswirkung lokaler Schwachstellen auf den mittels Durchbiegungsmessung 

bestimmten globalen Elastizitätsmodul gezeigt. Der aus den teilweise 

extrem variierenden lokal gemessenen Elastizitätsmoduln berechnete globale 

Elastizitätsmodul stimmt sehr gut mit dem gemessenen globalen Elastizitätsmodul 

überein. 

183 


S. AICHER, L. HÖFFLIN, W. BEHRENS 

Die experimentellen Untersuchungen wurden mit hochkant biegebeanspruchten 

Buchebrettern durchgeführt. Zuerst wurde die Längsverschiebungs- 

/Dehnungsmeßmethode exemplarisch an einem Brett verifiziert, welches stufenweise 

mit künstlichen Defekten (Löchern) versehen wurde. Für jeden Defektzustand 

wurden die lokalen und globalen Elastizitätsmoduln gemessen. In 

einem zweiten Schritt wird die Änderung des lokalen Elastizitätsmoduls und 

dessen hohe Korrelation mit dem Ort des Biegeversagens exemplarisch an vier 

Brettern einer größeren Versuchsreihe aufgezeigt. 

RESUME 

Cet article présente une méthode efficace permettant de déterminer le module 

d’élasticité local par une mesure couplée déplacement/déformation. D’autre 

part, on fait apparaître l’effet des sections localement faibles sur le module 

d’élasticité global déterminé par la flèche. Le module d’élasticité global obtenu 

par intégration des modules locaux mesurés, extrêmement variables, est en bon 

accord avec le module global mesuré. 

L’étude expérimentale a porté sur des panneaux fléchis à chant. La méthode 

couplée déplacement/déformation a été préalablement vérifiée sur un panneau 

présentant des défauts artificiels (trous). Pour chaque défaut, on détermine 

les modules local et global, et les différences sont discutées. Par la suite, la variation 

du module local et sa forte corrélation spatiale avec la résistance en 

flexion est mise en évidence sur 4 panneaux de hêtre extraits d’une campagne 

expérimentale plus importante 

KEYWORDS: local modulus of elasticity, global modulus of elasticity, stiffness 

variation, artificial defects, weak sections 

1 INTRODUCTION 

It is reported on a method for determination of the local modulus of elasticity 

(MOE) in bending tests with timber beams and respective results. In heterogeneous 

materials such as wood the modulus of elasticity can vary strongly 

along the length of the boards. Based on a positive stiffness - bending strength 

correlation, the footprints of locally low MOE values determine the strength 

class (or grade) of boards in grading machines based on the bending principle. 

Local MOE obviously depends strongly on the length of the board segment used 

184


for the MOE determination which in most cases is larger than a local weak area, 

mostly created by a knot or by sloping grain. 

A local MOE determined over a board segment length of 5 times the crosssectional 

depth as specified in EN 408 still represents an integral (constant) 

value over a considerable length and there may be strong local MOE deviations 

within that length. The stated averaging effect of concentrated zones of low 

MOE areas occurs in all bending type grading machines which bend at consecutive 

locations, as there are practical limits for the length of the span. This is the 

major reason for the moderate coefficient of correlation between bending 

strength and MOE. Interesting approaches on how to reconstruct the variation of 

the true MOE function from MOE data collected from a consecutively bent 

board in order to derive the true local MOEs based on Fourier transforms were 

proposed by Bechtel (1985) and Foschi (1987). 

Apart from strength grading the knowledge of the actual local MOE and of 

the associated local strength variation along the length of boards is very important 

for (stochastic) modelling of boards and glulam subdivided in unit cells of 

small length, i.a. Foschi and Barrett (1980), Ehlbeck et al. (1985), Isaksson 

(1999) and Serrano (2001). Hereby the length of the unit cell has a considerable 

modelling influence on load sharing in adjacent glulam lamellas. 

For modelling and calibrating the MOE variation along board length several 

approaches are known (i.a. Foschi and Barrett (1981), Ehlbeck et al. (1985), 

Kline at al. (1986) and Taylor (1991)). All models are based on a calibration vs. 

global (and partly local) modulus of elasticity necessitating extensive empiric 

data and leaving model dependent considerable uncertainties. 

The experimental determination of local MOEs which at first view seems 

to be a very simple task is demanding in case a bending method is applied and 

has limits concerning the smallness of the segment length. Reliable results below 

span to depth ratios of about 3 are questionable; limits were revealed by 

Kaas (1975) employing the so-termed “middle ordinate method”. The method is 

based on the assumption that short segments of a bent board approximate arcs of 

circles with varying radii. 

The work reported here was conducted in the frame of establishing a realistic 

empirical data basis for the variation of bending MOE and bending strength 

values along the length of beech wood boards bent about the major axis. It was 

185 



Table 1: Compilation of local compression and tension strains of test set A 

test 

A1 

A2 

A3 

compression 

and tension 

strain per 1kNm 

ε c 

ε t 

ε c 

ε t 

ε c 

ε t 

10 -5 

10 -5 

10 -5 

10 -5 

10 -5 

10 -5 

local strains in segments 1 to 6 

1 2 3 4 5 6 

-62.4 -60.4 -58.7 -64.9 -61.3 -60.2 

71.2 66.8 66.7 70.2 67.5 64.9 

-63.3 -59.5 -57.8 -66.7 -79.4 -62.9 

70.3 66.8 66.6 70.2 68.4 64.9 

-61.5 -57.7 -61.4 -66.7 -74.0 -61.1 

70.3 66.8 82.8 72.0 68.4 64.9 

Table 2: Compilation of results for local and global MOEs of test set A 

modulus of elasticity [N/mm 2 ] 

test measured local MOEs in segments 1 to 6 

- 

global 

measured 

MOE 

calculated MOE based 

on the measured local 

MOEs 

Eseg1 Eseg2 Eseg3 Eseg4 Eseg5 Eseg6 Eglob Eglob, calc 

N/mm 2 

N/mm 2 

N/mm 2 

N/mm 2 

N/mm 2 

N/mm 2 

N/mm 2 

N/mm 2 

A1 13495 14172 14381 13347 13989 14415 13919 13932 

A2 13494 14272 14493 13171 12197 14110 13340 13576 

A3 13686 14479 12502 13000 12660 14311 12879 13141 

15000 

14500 

14000 

13500 

13000 

12500 

12000 

1 2 3 4 5 6 

segment 

Eseg Eglob |εc| εt Fig. 3: Local strains, local and global MOEs of test A1 with board No. 1; no artificial 

defects 

192 

85 

80 

75 

70 

65 

60 

55 

strain [10 -5 /kNm]


15000 

14500 

14000 

13500 

13000 

12500 

12000 


1 2 3 4 5 6 

segment 

E seg E glob |ε c| ε t 

Fig. 4: Local strains, local and global MOEs of test A2 with board No. 1; artificial 

defects in the bending compression part of segment 5 


15000 

14500 

14000 

13500 

13000 

12500 

12000 

1 2 3 4 5 6 

segment 


Fig. 5: Local strains, local and global MOEs of test A3 with board No. 1; additional artificial 

defects in the bending tension part of segment 3 

193 

85 

80 

75 

70 

65 

60 

55 

85 

80 

75 

70 

65 

60 

55 

strain [10 -5 /kNm] 




5 RESULTS OF TEST SET B 

Table 3 specifies the measured local and global MOEs, the computed 

global MOEs and the location of the failure of four beech boards. The chosen 

examples are exemplary for 30 tests performed so far. Figures 6 to 9 give 

graphical representations of the results including the local strain variations. 

Figure 6 reveals the case of a nearly homogeneous board with no knots and 

no apparent grain deviation. Local and global MOEs show very little differences. 

Despite the small stiffness variations, the bending tension failure occurred 

in segment 3 with the lowest local MOE and with highest tension strain. The 

minimum local MOE differed only by 3% from the global MOE and only by 

1.3% from the next weakest segment 1. 

Figure 7 also depicts the strains and local MOE variations of a board without 

knots, but nevertheless with pronounced differences (maximally 18%) of 

local MOEs ranging from 12600 to 15400 N/mm 2 . The extreme deviations of 

the local MOEs vs. the global MOE are + 7.7% and – 11.8%. The strains of segments 

2 and 3 with lowest local MOEs show an interesting feature being that 

maximum tension and compression strain occur successively in segments 2 and 

3, indicating sloping grain. The specimen failed in bending tension at the transition 

of segments 2 and 3. 

Figure 8 relates to a board with a knot of 22 mm diameter and associated 

strong fibre deviations around the knot located in the upper bending compression 

part of segment 3. The very pronounced difference between the extreme 

local MOEs of 10360 and 14200 N/mm 2 was 27%; the extreme deviations of the 

local MOEs vs. Eglob were –17% and 13.6%. 

Table 3: Compilation of local and global MOEs of test set B 

test measured local MOEs in segments 1 to 6 

global 

measured 

MOE 

calculated MOE 

based on the 

measured local 

MOEs 

location of 

failure 

E seg1 E seg2 E seg3 E seg4 E seg5 E seg6 E glob E glob,calc segment 1) 

B1 13555 14140 13375 13746 13746 14138 13779 13713 3 

B2 15395 13687 12597 14032 14586 14212 14288 13703 2 - 3 

B3 13808 11696 10356 13435 14203 12911 12508 12222 2 - 3 

B4 17844 17843 16270 15825 15842 15366 16830 16333 4 

1) x - y means the intersection betw een tw o segments 

194


17000 

16000 

15000 

14000 

13000 

12000 


area of failure initiation 

1 2 3 4 5 6 

segment 


Fig. 6: Local strains, local and global MOEs of test B1 (board No. 416) 


16000 

15000 

14000 

13000 

12000 

11000 


1 2 3 4 5 6 

segment 



The specimen failed as sole specimen in the tests so far in bending compression 

at the transition of segments 2 and 3 with highest compression strains and lowest 

local MOE, respectively. 

195 

65 

60 

55 

50 

45 

40 

70 

65 

60 

55 

50 

45 






15000 

14000 

13000 

12000 

11000 

10000 


1 2 3 4 5 6 

segment 




19000 

18000 

17000 

16000 

15000 

14000 


1 2 3 4 5 6 

segment 



Figure 9 shows strains and MOEs of a board without knots and absolutely 

very “high” MOEs with a global MOE of 16830 N/mm 2 . Bending compression 

and tension strains are very similar. The specimen failed in bending tension in 

segment 4 with the second lowest MOE. However, local MOEs in segments 4, 5 

and 6 are very similar and deviate maximally by 2% from their respective mean. 

196 

100 

90 

80 

70 

60 

50 

65 

60 

55 

50 

45 

40 


strain [10 -5 /kNm]


Again, as in test set A, a very good agreement between the experimentally 

and computationally obtained global MOEs was observed. The deviation was in 

average 2.4% and maximally 3.7%. 

6 DISCUSSION 

The results show that the employed method is able to deliver local MOEs 

and therefore to reveal the MOE variation within a board. However, the measured 

local MOEs still do not represent the true MOEs of the zones with or without 

defects within the board. The measured MOE depends to a great extent on 

the gauge (segment) length, L, the length of the weak area and also on the relative 

differences of the stiffness within the gauge length. The smaller the chosen 

segment length, the smaller the difference between the measured and the “true” 

MOE. The employed gauge length of about 1.5 times the board depth seems to 

be in the size range of typical defect zones of the regarded wood species as the 

results show a good correlation between the minimum localized MOE value and 

location of bending failure. However, some improvement should still be obtained 

by a further reduction of gauge length L. 

7 CONCLUSIONS 

The presented results show that determination of the local modulus of elasticity 

in (edgewise) bending can be well performed by elongation/strain measurement 

at the bending tension and compression edges. 

The measured local MOEs and the experimental global MOE obtained from deflection 

measurement, are consistent. This results from the fact that the global 

MOE can be predicted by beam theory or FE analysis with an average error of 

about 2% on the basis of the local MOE of the segments, here chosen with a 

length of 200 mm. 

It was revealed that the locations of failure comply well with the locations of 

minimum MOE along beam length (the study so far comprised 30 beech 

boards). The presented method seems to enable the prediction of the type of 

bending failure either at the tension or compression edge. 

The data of the on-going study serve as a calibration basis for modelling of the 

variation of modulus of elasticity and bending strength along beech wood boards 

as input data for glued compound elements with edgewise bent lamellas. 

197 




The authors want to express sincere thanks to Patrick Castera, head of Laboratoire 

du Bois de Bordeaux (LRBB), for his repeated favour in performing 

the translation of the French abstract. 

REFERENCES 

BECHTEL, F.K. (1985): Beam stiffness as a function of point wise E, with application 

to machine stress rating. Proceedings International Symposium on 

Forest Products Research, CSIR, Pretoria, South Africa 

CORDER, S.E. (1965): Localized deflection related to bending strength of lumber. 

Second Symposium on the Non-destructive Testing of Wood, Washington 

State University, Pullman, WA, pp. 461 – 473 

EHLBECK, J., COLLING, F., GÖRLACHER, R. (1985): Einfluß keilgezinkter Lamellen 

auf die Biegefestigkeit von Brettschichtholzträgern. Entwicklung eines Rechenmodells. 

Holz Roh- Werkstoff, 43, pp. 333 – 337 

FOSCHI, R.O., BARRETT, D. (1980): Glued-laminated beam strength: A model. J. 

of the Structural Div., Vol. 106, No. ST8, pp. 1735 – 1754 

FOSCHI, R.O. (1987): A procedure for the determination of localized modulus of 

elasticity. Holz Roh- Werkstoff, 45, pp. 257 – 260 

ISAKSSON, T. (1999): Modeling the variability of bending strength in structural 

timber; length and load configuration effects. Report MBK-1015, Div. of 

Structural Eng., Institute of Technology, Lund, Sweden 

KASS, A.J. (1975): Middle ordinate method measures stiffness variation within 

pieces of lumber. Forest Products J., 25 (3), pp. 33 – 41 

KLINE, D.E., WOESTE, F.E., BENDTSEN, B.A. (1986): Stochastic model for modulus 

of elasticity of lumber. Wood and Fibre Science, 18 (2), pp. 228 - 238 

SERRANO, E. (2001): Mechanical performance and modeling of glulam. Manuscript 

for „Timber Engineering“, Edts. S. Thelandersson and H.J. Larsen, 

Wiley & Sons, in press 

TAYLOR, S.E., BENDER, D.A. (1991): Stochastic model for localized tensile 

strength and modulus of elasticity in lumber. Wood and Fibre Science, 

23(4), pp. 501 - 519 

198

Transient temperature evolution in glulam with hidden and non-hidden glued-in steel rods 

TRANSIENT TEMPERATURE EVOLUTION IN GLULAM WITH 

HIDDEN AND NON-HIDDEN GLUED-IN STEEL RODS 

TRANSIENTE TEMPERATURENTWICKLUNG IN BRETTSCHICHT- 

HOLZ MIT VERDECKT UND NICHT VERDECKT EINGEKLEBTEN 

STAHLSTANGEN 

EVOLUTION TRANSITOIRE DE LA TEMPERATURE DANS DU 

LAMELLE COLLE COMPORTANT DES GOUJONS COLLES EN 

ACIER, APPARENTS OU NON 

Simon Aicher, Dirk Kalka, Ralf Scherer 

SUMMARY 

A recently terminated European research project on glued-in steel rods in 

timber structures (GIROD) – with participation of Timber Department of Otto- 

Graf-Institute – revealed a strong strength reducing influence of elevated temperatures, 

not expected to that extent. This affects especially the duration of load 

behavior of the connections. The maximum temperature level acting in service 

on the glued-in rod connections thus sets performance requirements on the shear 

modulus-temperature relationship and on the glass transition temperature of appropriate 

adhesives. 

Today’s prevailing intuitive conviction of practitioners is, that rods bonded 

hidden in the interior of glulam cross-sections experience, due to the low thermal 

conductivity and specific heat of wood, considerable lower temperatures 

compared to ambient climate. The paper gives some experimental and computational 

results proving, depending on cross-sectional width, only a moderate reduction 

of peak temperatures combined with a pronounced phase shift vs. ambient 

temperature varying roughly sinusoidally during a day. 


Ein kürzlich abgeschlossenes Europäisches Forschungsvorhaben betreffend 

in Holz eingeklebter Stahlstangen (GIROD) – mit Beteiligung des Fachbereichs 

Holz des Otto-Graf-Instituts – zeigte einen in dieser Ausprägung nicht erwarteten 

großen festigkeitsmindernden Einfluss erhöhter Temperaturen. Dies beeinflusst 

insbesondere auch das Zeitstandverhalten der Verbindungen. Das maxi- 

199 


S. AICHER, D. KALKA, R. SCHERER 

male Temperaturniveau, das im Gebrauchszustand auf Verbindungen mit eingeklebten 

Stangen einwirkt, definiert somit Leistungsanforderungen an die 

Schubmodul-Temperaturbeziehung und an die Glasübergangstemperatur geeigneter 

Klebstoffe. 

Die heute in der Praxis vorherrschende Meinung ist, dass verdeckt in das 

Innere eines Brettschichtholzquerschnitts eingeklebte Stangen infolge der niedrigen 

Wärmeleitfähigkeit und Wärmekapazität von Holz beträchtlich niedrigeren 

Temperaturen im Vergleich zum einwirkenden Umgebungsklima ausgesetzt 

sind. Der Aufsatz berichtet über einige experimentelle und rechnerische Ergebnisse, 

die belegen, dass abhängig von der Querschnittsdicke lediglich eine 

schwache Reduzierung der Spitzentemperaturen verbunden mit einer ausgeprägten 

Phasenverschiebung gegenüber den Außentemperaturen, die näherungsweise 

sinusförmig über den Tag variieren, vorliegt. 

RESUME 

Un projet de recherche Européen portant sur les goujons collés en acier 

dans les structures en bois (GIROD) récemment achevé – auquel participait le 

département bois de l’Otto-Graf Institute – a mis en évidence un effet négatif 

marqué de températures élevées sur la résistance, qui affecte principalement la 

durée de vie des joints. La température maximale agissant sur les joints en service 

impose donc des exigences de performance sur la relation température – 

module de cisaillement et la température de transition vitreuse des adhésifs appropriés. 

La conviction intuitive des praticiens aujourd’hui est de penser que les goujons 

collés cachés à l’intérieur des sections de lamellé collé sont soumis, du fait 

de la faible conductivité thermique et chaleur spécifique du bois, à des températures 

considérablement plus faibles que celles du climat ambiant. Cet article présente 

des résultats expérimentaux et numériques montrant, selon la largeur de la 

section, une faible réduction seulement des températures de pic combinée à une 

transition de phase prononcée, par rapport à la température ambiante qui varie 

grossièrement de manière sinusoïdale au cours d’une journée. 

KEYWORDS: glued-in rods, glulam, elevated temperatures, transient temperature 

evolution 

200



Today’s prevailing conviction of practitioners is, that steel rods bonded 

hidden in the interior of glulam cross-sections experience, due to the low thermal 

conductivity and specific heat of wood, considerable lower temperatures 

compared to ambient climate. A European research project on glued-in rods in 

timber structures (GIROD) revealed an unexpected strong strength reducing influence 

of elevated temperatures in duration of load tests with connections 

bonded by epoxy and polyurethane adhesives [BENGTSSON, JOHANSSON, 2002; 

AICHER, 2002]. 

In the performed tests the temperature increase was applied after mechanical 

loading thus suppressing positive post-curing effects. The experiments revealed 

clearly the crucial importance of a sufficiently high glass transition temperature. 

Performance requirements on temperature stability – especially shear 

modulus-temperature relationships and/or glass transition temperature – have to 

be set in view of realistic temperature loading scenarios. Eventually the temperature 

loading should also be considered in probabilistic manner, in case a specific 

adhesive shows high post-curing potential. 

The reported experimental investigations were performed in first instance 

to substantiate the GIROD results. In addition thereto a major point of interest 

was the transient temperature evolution in the timber-bond line interface. 

2. TEST PROGRAM 

In order to verify the different temperature-strength behavior of glued-in 

steel rods either protruding or fully hidden in the wood, two types of specimens 

shown in Fig. 1 were investigated. The performed experiments concerned the 

strength verification at variable temperature and static long term loads and furthermore 

the temperature evolution in the bond lines. This paper reports on the 

temperature evolution. 

The temperatures in the bond line were measured with thermo-elements 

consisting of copper/constantan (Cu/Cu-Ni) wires. In order to measure the temperatures 

in the bond line with little interfering influences of leakages to ambient 

climate, the application of the thermo-element wires to the bond line was performed 

as following: first an oversized specimen was sawn up lengthwise with a 

saw blade thickness of 2 mm. Then the two parts were clamped together and a 

hole with a diameter of 13 mm for the glued-in rod was drilled. The thermo- 

201 



wires were glued into small notches as shown in Fig. 2a. The end or actual 

measuring point of the thermo-wire was flush with the surface of the drilled 

hole. Then the two parts of the specimen were glued together again. In a second 

step the steel rod (Ø 12 mm) was glued into the wooden piece (see Fig. 2b). All 

gluing were performed with a special epoxy adhesive. 

thermoelements 

40 

40 

40 

40 

15 

5 

T1 

T2 

T3 

T4 

T5 

180 

240 

180 

glued-in test 

steel rod 

M12 

600 

115 

115 

specimen 

part B 

thermoelements 

40 

40 

40 

40 

specimen 

part A 

support rod 

M24 

a) b) 

T1 

T2 

T3 

T4 

T5 

Fig. 1a,b: Geometry and schematic test set-up of 

a) specimen No. I with protruding steel rod and 

b) specimen No. II with hidden steel rod 

15 

5 

support rod 

M24 

180 

240 

180 

600 

support rod 

M24 

glued-in test 

steel rod 

M12 

600 

115 

115 

insulation 

tape 

In order to obtain a specimen with a fully hidden rod which could be subjected 

to temperature and mechanical loads, specimen No. II, shown in Fig. 1b, 

202


was used. First, specimen part A was manufactured as specified above and after 

curing of the adhesive, part A incorporating the protruding test rod was glued 

into the rod hole of specimen part B. In order to enforce exclusively load transfer 

between the specimen parts A and B via the glued-in rods, a Teflon sheet 

with a thickness of 0.5 mm was inserted between the two parts of the specimen 

(see Fig. 2c) . The surrounding edge of 10 mm width and 2 mm depth of the two 

specimen parts was sealed with an elastic insulation tape compressed to 0.5 mm 

(see Fig. 1b and 2d). 

a) b) 

c) d) 

Fig. 2a-d: Views of the specimens No. I (a,b) and No. II (c,d) 

203 



3. TEMPERATURE LOADING 

As in the GIROD project a cyclic sinusoidal variation of warm and dry 

climate was applied [AICHER ET AL., 2002]. Contrary to GIROD, where a full 

temperature cycle consisted of 8 hours, now a practically relevant cycle length 

of 24 hours was chosen. A sinusoidal variation of temperature within a time 

span of 24 hours represents a very good approximation of daily temperature 

courses. This is shown exemplarily in Fig. 3 with recorded temperature data 

(sheltered outdoor, well ventilated shed in Stuttgart) for a period of three successive 

days. 

temperature T [°C] 

30 

28 

26 

24 

22 

20 

18 

16 

sinusoidal approx. 

measured 

14 

28.7.02 29.7.02 30.7.02 31.7.02 

Fig. 3: Course of temperature (sheltered outdoor conditions) of typical summer days and 

sinusoidal approximation of the temperature 

As in the GIROD project the minimum and maximum set temperatures 

were chosen as 25°C and 55°C, resulting in a peak-to-peak temperature amplitude 

of 30 K. These temperature boundaries might be regarded as an upper, yet 

realistic temperature range, which can occur under a dark, little ventilated roof 

in very warm summers in the Southern part of Europe. The course of the applied 

temperature and of the relative humidity is given in Fig. 4. The controlling of the 

relative humidity was limited, due to technical restrictions of the climate chamber, 

to 45% RH during a time of 6.5 h of a full cycle of 24 h, as shown in Fig. 4. 

204 

time


The actual temperatures obtained in the climate chamber showed a minimum 

and maximum of 24.9°C and 54.7°C, respectively, with a peak-to-peak 

temperature amplitude of 29.8 K. The relative humidity roughly ranged from 5 

to 50 %, exceptionally temporary up to 67 % for about 0.5 hours. 

temperature T [°C] / 

relative humidity RH [%] 

70 

60 

50 

40 

30 

20 

10 

0 

T 

RH 

0 12 24 36 48 60 72 84 96 108 120 

time t [h] 

Fig. 4: Course of the applied ambient temperature and relative humidity variation in the 

climate chamber 

4. NUMERICAL INVESTIGATIONS 

In an early paper the evolution of temperature in a specimen with a gluedin 

rod protruding into ambient air was investigated numerically and experimentally 

[AICHER ET AL, 1998], taking into account the timber, the adhesive layer 

and the steel. An additional fourth layer, representing a steel/adhesive interface, 

was introduced in order to account for the problem that the used FE-code does 

not enable the specification of contact conductance of inner surfaces. By means 

of the interface layer a good agreement of measured and calculated transient 

temperatures was obtained. 

205 



In this paper the preliminary numerical study is exclusively related to 

specimen No. II with the hidden rod. In a first crude approximation the inner 

steel rod was omitted, so only the heat transfer through a quadratic block of timber 

was regarded in a 2 dimensional analysis. The cross-sectional dimensions of 

the timber, a = 115 mm, were reduced by the hole diameter of 13 mm (rod 

diameter + 1 mm) to 102 mm. 

D 

In the calculations a constant thermal diffusivity perpendicular to grain of 

k mm² 

700 

C 

h 

P ρ 

was employed. Hereby k , CP and ρ were assumed as 

k 

0. 

13 

W 

m K 

thermal conductivity perpendicular to fiber 

acc. to DIN 4108, part 4 

kJ 

C P 1. 

6 

specific heat [BATZER, 1985] 

kg K 

ρ 

kg 

420 

m³ 

mass density of glulam at dry status of about 

u = 7 % 

The convection heat transfer coefficient h was chosen as a fitting parameter 

in the range of 10 to 20 W/(m²·K). Literature data for forced convection of gas 

media vary roughly between 10 and 100 W/(m²·K); a value of 25 W/(m²·K) is 

assumed for convection at exterior walls in DIN 4108, part 4. 

5. TEMPERATURE EVOLUTION IN CYCLIC CLIMATE 

Figs. 5a,b show the temperature evolution of both specimen types I and II 

at different thermo-element positions. The evolution of the ambient temperature 

in the climate chamber is given, too. For a better visualization of phase shift and 

differences in amplitudes Figs. 6a and b show the temperature evolution at a cycle 

length of 24 hours; additionally finite element computed temperature evolutions 

based on the revealed approach are specified in case of specimen type II 

with a hidden rod. 

206


a) 

[°C] 

temperature T 

b) 

60 

55 

50 

45 

40 

35 

30 

25 

20 

60 

55 

50 

45 

40 

35 

30 

25 

20 


ambient climate (climate chamber) 

0 6 12 18 24 30 36 42 48 54 60 66 72 

T1 - T5 

measured 

temperatures 

time t [h] 

T1 

T2 

T3 - T5 

ambient climate (climate chamber) finite element calc. 

0 6 12 18 24 30 36 42 48 54 60 66 72 

time t [h] 

calc_1 

calc_2 

calc_3 

Fig. 5a,b: Temperature evolution over 3 days at the positions of the thermo-elements 

a) specimen No. I with protruding steel rod 


207 

T1 T0 

T2 T1 

T3 T2 

T4 T3 

T4 

T5 

part A 

T5 T 

T4 T 

T3 

T2 

T1 

T 

T 

T 

part B 




a) 

[°C] 

ure T 

temperat 

b) 

60 

55 

50 

45 

40 

35 

30 

25 

ambient climate 

(climate chamber) 

T1 

T2 

T3 - T5 

20 

24 27 30 33 36 39 42 45 48 

60 

55 

50 

45 

40 

35 

30 

25 



time t [h] 

measured 

temperatures 

20 

24 27 30 33 36 39 42 45 48 

time t [h] 

T1 - T5 

finite element calc. 

calc_1 

calc_2 

calc_3 

Fig. 6a,b: Temperature evolution over 1 day at a cycle length of 24 hours 

a) specimen No. I with protruding steel rod 


208 

T1 T0 

T2 T1 

T3 T2 

T4 T3 

T4 

T5 

part A 

T5 

T4 

T3 

T2 

T1 

part B


The differences between the two specimen types are rather small. Purely 

qualitatively the temperature in the wood-bond line interface of specimen No. II 

shows slightly decreased amplitudes and a slightly more pronounced phase shift 

vs. ambient climate when compared to specimen No. I with the protruding rod. 

Quantitatively the results are specified in Tab. 1. 

Tab. 1: Temperature evolution in the wood-adhesive interface of specimens No. I and 

No. II at thermo-element positions T1 and T5 

ambient 

climate 

maximum 

temperature 

Tmax 

experimental results: 

specimen 

No. I 

(protruding 

rod) 

specimen 

No. II 

(hidden 

rod) 

thermo-element T1 

(“protruding” end of steel rod) 

minimum 

temperature 

Tmin 

peak-topeak 

amplitude 

∆T 

phase 

shift 

∆t 

maximum 

temperature 

Tmax 

thermo-element T5 

(embedded end of steel rod) 

minimum 

temperature 

Tmin 

peak-topeak 

amplitude 

∆T 

phase 

shift 

[°C] [°C] [K] [h] [°C] [°C] [K] [h] 

54.7 24.9 29.8 - 54.7 24.9 29.8 - 

53.4 28.1 25.3 1.7 52.6 28.2 24.4 2.4 

51.2 28.7 22.5 3.3 51.6 28.8 22.8 3.2 

It can be seen that the maximum temperatures at the embedded ends of the 

rods (thermo-element T5 for specimens No. I and No. II) differ only very little 

by about 1°C. The reduction of maximum temperature vs. ambient climate in 

case of specimen No. II (hidden rod) was only 3 K. The phase shift between the 

maximum temperature of the ambient climate and the maximum of recorded 

temperatures was 2.4 and 3.2 hours in case of specimens No. I and No. II, respectively. 

In case of specimen No. II no difference of temperature amplitudes, peak 

temperatures and phase shifts between thermo-element T1 close to the sealed 

joint of both specimen parts A and B as compared to the embedded end (thermoelement 

T5) was observed. In case of leakages at the sealing of the joint of the 

209 

∆t 



parts A and B a different result should be obtained. This is substantiated by the 

calculations. 

Table 2 represents the results of the rough finite element calculation compared 

to the experimental results of specimen type II and the applied ambient 

climate. It can be seen that either the maximum and minimum temperatures Tmax 

and Tmin (result calc_1) or the phase shift ∆t (result calc_3) of the measured experimental 

results can be fitted by tuning of the convection heat transfer coefficient 

h. A roughly acceptable approximation of both , the temperatures and the 

phase shift, is obtained with a convection heat transfer coefficient of h = 15 

W/(m²·K). This number is within the plausible range. 

Tab. 2: Maximum temperatures and phase shift for specimen No. II according to experimental 

test results and finite element calculation 

ambient 

climate 

experimental 

results 

2D finite 

element 

calculation 

calculation 

result 

convection 

heat transfer 

coefficient 

h 

[W/(m²·K)] 

maximum 

temperature 

Tmax 

[°C] 

minimum 

temperature 

Tmin 

[°C] 

peak-to-peak 

amplitude 

∆T = 

Tmin-Tmax 

[K] 

- - 54.7 24.9 29.8 - 

phase 

shift 

- - 51.2 28.7 22.5 3.3 

calc_1 10 51.4 28.6 22.8 4.2 

calc_2 15 52.6 27.4 25.2 3.6 

calc_3 20 53.2 26.8 26.4 3.3 

6. INFLUENCE OF TIMBER THICKNESS 

The presented test results and hereby the small damping of the temperature 

maxima is obviously related to the cross-sectional dimensions of the specimens 

(quadratic cross-section of 115 mm · 115 mm). In order to verify the influence 

of an increased timber thickness some more calculations similar to those outlined 

in chapter 4 were performed. The convection heat transfer coefficient was 

chosen as h = 15 W/(m²·K) being the value which forwarded a reasonable good 

agreement of the simplified analysis with the hidden rod specimen No. II. The 

imposed temperature varies again sinusoidally between 25 and 55°C with a 

phase length of 24 hours. 

210 

∆t 

[h]


Fig. 7 gives the temperature courses for square cross-sectional dimensions 

with thicknesses of a = 50, 102, 150 and 200 mm. The value a = 102 mm relates 

to the discussed results of specimen No. II, whereby the reduction of the real 

thickness of 115 mm to 102 mm is bound to the simplified approach of omitted 

rod cross-section. The graph shows the qualitatively somewhat trivial result of a 

decreasing temperature amplitude and an increasing phase shift with growing 

cross-sectional dimensions. 


60 

55 

50 

45 

40 

35 

30 

25 



measured 

temperatures 

T1 - T5 

finite element calc. 

a = 50, 102, 150, 200 mm 

20 

24 27 30 33 36 39 42 45 48 

time t [h] 

Fig. 7: Temperature evolution at a cycle length of 24 hours for finite element calculations 

with varying timber thicknesses compared to experimental results (specimen No. II 

with hidden rod) 

A quantitative summary of the results is given in Tab. 3. In detail the 

changes of minimum and maximum temperature, the peak-to-peak amplitude 

∆T = Tmax-Tmin and the phase shift ∆t are specified. It can be seen that the reduction 

of the maximum temperature in the regarded range of cross-sectional dimensions 

is rather moderate. For a medium sized glulam thickness of 150 mm 

the maximum value is still rather close to 50°C, which marks the limit of type I 

adhesives acc. to EN 301. 

The quantitative results of the very rough approximation of the problem 

shall be regarded with a more refined modeling considering the true build-ups. 

211 

a 

T5 

T4 

T3 

T2 

T1 



Tab. 3: Extreme temperatures, temperature differences and shifts of phase depending on 

timber thickness according to a simplified calculation 

ambient 

climate 

experimental 

results 

2D finite 

element 

calculation 

calculation 

result 


crosssectional 

thickness 

a 

[mm] 

maximum 

temperature 

Tmax 

[°C] 

minimum 

temperature 

Tmin 

[°C] 

peak-to-peak 

amplitude 

∆T = Tmin-Tmax 

[K] 

phase shift 

- - 54.7 24.9 29.8 - 

- - 51.2 28.7 22.5 3.3 

calc_1 50 54.0 26.0 28.0 2.5 

calc_2 102 52.6 27.4 25.2 3.6 

calc_3 150 48.7 31.3 17.4 5.7 

calc_4 200 46.3 33.6 12.7 7.2 

The performed experiments on transient temperatures in glue-line/wood 

interfaces of steel rods bonded into glulam and subjected to cyclically varying 

ambient climate revealed 

• relatively low damping of maximum temperatures for a cross-sectional 

thickness of 115 mm, 

• pronounced phase shifts and 

• only minor differences between the cases of protruding or hidden rods. 

The results were extrapolated to different cross-sectional thicknesses by 

means of numerical calculations with a simplified model. The calculation results 

yielded rather moderate damping within the typical range of glulam thicknesses 

up to 200 mm. Roughly it can be concluded that the maximum ambient temperature 

level acting in service on the glued-in rod connections sets the performance 

requirements on the shear modulus/temperature relationship resp. on the glass 

transition temperature of appropriate adhesives. 


The authors are cordially indebted to Dr. Patrick Castera, Head of Laboratoire 

du Rheologie du Bois Bordeaux (LRBB), for performing the french translation. 

212 

∆t 

[h]

REFERENCES 


AICHER, S. (2002): Duration of load tests on full-sized glued-in rod specimens. 

GIROD_WP5: Technical Report for work package 5, Research Report, 

Otto-Graf-Institute, University of Stuttgart. 

AICHER, S.; KALKA, D.; HÖFFLIN, L. (2001): Duration of load tests on fullsized 

glued-in rod specimens. GIROD_WP5: Technical Report for work 

package 5, workpart by FMPA. Research Report, Otto-Graf-Institute, 

University of Stuttgart. 

AICHER, S.; WOLF, M.; DILL-LANGER, G. (1998): Heat flow in a glulam joist 

with a glued-in steel rod subjected to variable ambient temperature. 

Otto-Graf-Journal Vol. 9, pp. 185-204 , Otto-Graf-Institute, University of 

Stuttgart. 

BATZER, H. (1985): Polymere Werkstoffe. Volume I, Georg Thieme Verlag 

Stuttgart. New York. 

BENGTSSON, C.; JOHANSSON, C.-J. (2002): GIROD – Glued in rods for timber 

structures. SP Report 2002:26. SP Swedish National Testing and Research 

Institute. 

213 



214

MODELLING OF CONCRETE HYDRATION 

MODELLIERUNG DER BETON HYDRATATION 

MODELISATION DE L'HYDRATATION DU BETON 

Sven Mönnig 

SUMMARY 

Modelling of concrete hydration 

DuCOM is a finite element program which can show the hydration of 

concrete with any concrete mixtures, to any given time step and different 

environmental conditions. Comparing calculated temperature distribution, 

hydration and heat growth rates with measurements a high accuracy was proven. 


DuCOM ist ein FEM-Programm, welches die Hydratation von Beton mit 

beliebigen Betonmischungen, zu beliebigen Zeitschritten und verschiedenen 

Umgebungsbedingungen darstellen kann. Bei dem Vergleich von errechneten 

Temperaturverläufen, Hydratationskurven und Wärmezuwachsraten mit 

gemessenen Laborwerten, zeigt sich eine hohe Genauigkeit von DuCOM. 

RÉSUMÉ 

DuCOM est un programme d'éléments finis capable de décrire l'hydratation 

de bétons de compositions arbitraires, à des intervalles arbitraires et pour 

différentes conditions d'environnement. La comparaison des gradients de 

température, des courbes d'hydratation et des taux de chaleur calculés avec les 

valeurs mesurées en laboratoire indiquent une précision élevée de DuCOM. 

KEYWORDS: DuCOM, heat growth rate, concrete hydration 


DuCOM [1] is a program working with concrete finite elements. It is able 

to deliver a linear description of the hydration of concrete. It provides solutions 

for pore pressure and temperature at each node of each element for given time 

steps and environmental condition, i.e. relative humidity and temperature. 

215 


S. MÖNNIG 

Results of porosity, the degree of hydration for every clinker, shrinkage and 

strength are obtained, too. By implementing DuCOM, a program developed by 

the University of Tokyo, into MASA [2] the simulation of the influence of 

hydration on the bearing capacity is possible. To estimate the computational 

accuracy of DuCOM extended calculations were compared with publications of 

test results. 

2. THEORETICAL PRINCIPALS OF DUCOM 

Physical processes like humidity and vapour transport, the hydration of 

concrete and the development of pore structure are integrated over the volume of 

a standard reference element. The transport behaviour is simulated on a macro 

scale. Hydration is simulated by a multi component system which includes the 

heat development and the amount of available water. The heat development is 

dependent on the amount of free water. Size and structure of the pores are 

dependent on the degree of hydration. The pore structure influences the transport 

behaviour inside the concrete. All the single processes are dynamically linked 

and dependent on each other as figure 1 should point up. 

FEMAP 

Visual Basic program 

DuCOM 

model for hydration 

pore structure 

pore pressure 

convergence 

T > T soll 

output of data 

VB program 

provides input data: 

nodes and elements; geometry of model 

addional input data: 

mixture; environmental conditions 

Program rewrites DuCOM output 

format into neutral ASCII files 

FEMAP display of results 

DuCOM developed by University of Tokyo 

output data: 

temperature; hydration degree 

dispersal of porosity 

relative humidity; dispersal of 

moisture; pore pressure 

Figure 1: application flow 

216


For further information it is recommended to read “Modelling of concrete 

performance” [1]. 

3. LIMITATIONS OF DUCOM 

There are some constrictions of DuCOM that should be mentioned. The 

pore structure is simulated by a consistent distribution of average sized grains. 

The size is dependent on the amount of cement, fly ash and blast furnace slag. 

The distance between the grains is based on Blaine values and the size of the 

grains. Pores are considered to be cylindrically shaped. For the calculation of the 

hydration the gel and capillary pores are treated as one type. 

The assumptions for the moisture transport are non deformable and 

isothermal material behaviour. Furthermore it is assumed that the total mass of 

vapour can be neglected compared to the total amount of water. Gas pressure 

within the material is constant and equals the air pressure. Liquid transport is 

performed with constant velocity. Thermal effects are negligibly small. These 

assumptions are based on a representative volumetrically element. All 

calculations refer to this element. 

4. IMPLEMENTATION OF DUCOM 

With FEMAP as input and output program it is possible to use a common 

used program for the visualization of the models. The output file from FEMAP 

is written in an ASCII format which is translated into the input file for DuCOM. 

For this transformation a Visual Basic (VB) program was developed by the 

IWB. While translating the file, the program asks for additional input data. 

Necessary input information are the time period of the analysis, mixture, Blaine 

values, temperature of the concrete mixture, temperature and relative humidity 

of the environment. After the end of the calculation another VB program will 

translate the result files of DuCOM into an ASCII format which can be read by 

FEMAP. The results can be graphically presented and furthermore they can be 

read from MASA and used for continuous analysis of the structure. 

5. RESULTS OF SIMULATIONS AND COMPARISON WITH TEST 

RESULTS 

All calculations were based on a 20×20×20 cm 3 cube. The resulting values 

of the curves were taken at nodes arranged along a line by the centre of a cube. 

217 


S. MÖNNIG 

OPC was simulated with these fractions of clinker: 

C3S 47,2 % 

C2S 27,0 % 

C3A 10,4 % 

C4(A,F) 9,4 % 

The model of the concrete cube has been scaled by the input program to the 

size desired by the user. Figure 2 shows the cube before scaling. 

Y 

10. 

9. 

element 27 

Element 27 

node Knoten 121 

X 

10. 

9. 

8. 

7. 

6. 

5. 

4. 

3. 

2. 

8. 

7. 

6. 

1. 

5. 

4. 

3. 

2. 0. 

1. 

0. 

Temperature distribution in a cube 

Knoten node 96 96 

Knoten 71 node 71 

node Knoten 46 

Knoten node 21 21 

Figure 2: Model of a 20×20×20 cm 3 cube 

Of interest was the influence of the environment on the development of 

heat inside the cube. Simulations with an adiabatic system have been performed 

as well as calculations with one, two, three, four and five sides opened to the 

environment. The mix temperature was 20°C. The environmental conditions 

have been assumed constant with 15°C and 100% relative humidity. The 

concrete mix contained 375 kg/m 3 cement and 1885 kg/m 3 aggregates. 

218

0.00 0.01 0.10 1.00 10.00 100.00 

Figure 3: Temperature distribution inside a cube 

Degree of Hydration in dependence on the water/cement ratio 


The proportion has been changed to a mixture with very high cement 

content. It contained 836.3 kg/m 3 cement and 1032 kg/m 3 aggregates. These 

fractions have been chosen to minimize the influence of the aggregates on the 

water diffusion and the hydration. The reference mixture had a water/cement 

ratio of 0.4 but the same cement and aggregate content as the others. This 

mixture has been calculated with five sides open to the environment which had a 

constant temperature of 15°C and a relative humidity of 100 %. 

W/C 0.4 Five Open 

Sides 

W/C 0.20 Adiabatic 



Hydration Degree [%] 

Progress of hydration 

0,0 0.0 

0 4 8 12 16 20 24 

Hours [h] 

Figure 4: Degree of hydration in dependence on the water/cement ratio 

219 

1,0 1.0 

0,9 0.9 

0,8 0.8 

0,7 0.7 

0,6 0.6 

0,5 0.5 

0,4 0.4 

0,3 0.3 

0,2 0.2 

0,1 0.1 


S. MÖNNIG 

Heat growth rate of clinker 

DuCOM provides the overall generated heat for each clinker at each time 

step. By subtracting the accumulated heat at one time step from the previous one 

it was possible to calculate the heat growth rate. The curves presented in figure 5 

have been interpolated with Excel to abrade them. DuCOM calculated 2500 time 

steps to reach 24 hours. 

Heat Growth Rate [kcal/kg] 

0,30 0.30 

0,25 0.25 

0,20 0.20 

0,15 0.15 

0,10 0.10 

0,05 0.05 

0.00 0,00 

0 4 8 12 16 20 24 

6. DISCUSSION 

Time [h] 

Figure 5: Heat growth rate of clinker 

C3A 

C3S 

C4AF 

C2S 

Total Heat 

The maximum temperature of the hydration as shown in figure 3 does 

reach the extent as expected. Different methods of gaining an approximated 

value do provide similar results, e.g. the approximation formula for adiabatic 

heat growth as given in [3] provides likewise results. The results presented in 

figure 5 follow the expected curves. The influence of the low temperature of the 

environment on the rate of hydration is reasonable, too. The curves in figure 5 

show the same behaviour of the clinker as it was expected due to the specific 

enthalpy of each clinker. 

7. SUMMARY OF RESULTS 

DuCOM proved to be very reliable being used for the simulation of 

hydration of ordinary Portland cements and their mixtures. Temperature 

development and hydration degree are corresponding with measured values 

given in the literature. More calculations and experiments should be performed 

to estimate the accuracy of calculated strength and shrinkage. 

220

REFERENCES 


[1] Maekawa, K.; Chaube R., Kishi, T.: Modelling of concrete performance, 

E&FN Spon, London, 1999 

[2] Ožbolt, J.: MASA – Macroscopic Space Analysis. Internal Report, Institut 

für Werkstoffe im Bauwesen, Universität Stuttgart, 1998 

[3] Zement Taschenbuch 2000, Verlag Bau+Technik, Düsseldorf, 2000 

221

CORROSION INDUCED FAILURES OF PRESTRESSING STEEL ...

You also want an ePaper? Increase the reach of your titles

Delete template?

Save as template?