bolted connections with hot dip galvanized steel members with

BOLTED CONNECTIONS WITH HOT DIP GALVANIZED STEEL MEMBERS WITH

PUNCHED HOLES

G. Valtinat, Technical University Hamburg-Harburg, Germany H. Huhn, Technical University Hamburg-Harburg, Germany

ABSTRACT Bolted connections of hot dip galvanized steel members with punched holes are usually taken for mast and tower constructions. These structures are often loaded by wind, so it makes good sense to examine the fatigue behaviour of the connections. One of the findings described in this paper is that the punching of the holes and the galvanizing process have a negative influence on fatigue behaviour. This poses the question: how can the fatigue resistance of bolted steel connections of galvanized steel members with punched holes be improved? The main idea is to use preloaded bolts. A preload perpendicular to the surface of the members protects the area around the hole by reducing the notch stresses. This means the structures can be strengthened and their lifetime will be prolonged.

INTRODUCTION In a large of test programme we investigated the complete behaviour of bolted connections of steel members in various applications, notably on masts and towers for electric transmission lines, antenna towers and lattice towers for wind turbines. These connections are usually made of hot dip galvanized steel members. Galvanizing is the best corrosion protection for this type of construction because the cleaning of heavily corroded masts later on, such as by shot blasting and repainting, is extremely expensive and time consuming. So the hot dip galvanizing of all steel members is advantageous. But considering the load carrying behaviour of the bolted connections, the sensitivity of the load displacements and the fatigue resistance, there are also some disadvantages brought about by hot dip galvanizing. Another important influence on these properties is the production of the bolt holes themselves. For the profiles used, such as angle profiles, it is much cheaper to punch the holes than to drill them. Punching produces a very heavy impact on the steel until a severe material hardening occurs around the holes. The ductility is reduced tremendously, the ultimate tensile strength is usually affected and cracks may be initiated. The area affected extends about 2 to 3 mm around the holes. This is the reason why reaming of the holes is often required with thicker plate material. Hot dip galvanizing (with its high temperature) that takes place after punching may also promote the aging of the material. All these influences have a negative effect the fatigue behaviour of the connections. STATIC TESTS Before we investigated this special problem it was necessary to gain knowledge on the influence of punching and hot dip galvanizing on the load carrying behaviour of bolted

Connections in Steel Structures V - Amsterdam - June 3-4, 2004 297

connections under static loads. We had several research programmes (1, 2) on the relations of load displacement compared with those of bolted connections with both galvanized and non galvanized members with drilled holes. Figure 1 shows the load displacement lines of two typical double sided lap connections, which are designed to break by ovalisation and bearing of the holes, which were the weakest part of the connection. We see that both connections reach nearly the same ultimate capacity (about 300 kN), but the connection with drilled holes (dashed line) has a much higher plastic displacement than that one with punched holes. Nevertheless, a connection which reaches 6 mm displacement due to hole ovalisation has, in our opinion, enough ductility for static loading.

Figure 1. Load-displacement curves of bolted double sided lap connections with drilled and punched holes – rupture by ovalisation and bearing.

Figures 2 and 3 show the difference of the rupture modes of two connections which were designed to break by net section rupture. In figure 2 we see the net section of a member with a drilled hole; only one crack runs through the section with a high reduction in plastic thickness reduction due to Poisson's ratio. By contrast, figure 3 shows the net section of a member with a punched hole. Many initial cracks start from the hole wall, and one of them finally runs through the net section but without any reduction in plastic thickness. This confirms the previous assumption that the ductility is very much reduced.

Figure 2. Net section rupture of a member Figure 3. Net section rupture of a member with a drilled hole in a static with a punched hole in a static short term test. short term test.

298 Connections in Steel Structures V - Amsterdam - June 3-4, 2004

MATERIALS TESTING The structure of the material at the edge of a punched hole was examined (3). Microscopic photos were taken at different levels of the thickness of the plate. They show how much the material along the hole wall changes its grain structure when the punch is driven through it. Just after the entrance of the punch a special zone at the upper part of the plate is formed, figure 4a. In this first zone the material flows in the cutting direction of the punch from the edge into the hole. The plastic flow passes over to the second zone in the middle of the plate thickness, figure 4b. Then shear cutting takes place, and is followed by the third zone where the punch leaves the plate, shown in figure 4c. Here a rougher surface is built which is widened conically.

Figure 4. Micro-section of the material structure at the edge of a punched hole (calibration line 0.1 mm). a) Zone at the upper side when punch enters the plate b) Section in middle of the member c) Zone at the lower side when punch leaves the plate

In addition, the distribution of the hardness in different sections parallel to the surface at intervals of 0.5 mm has been worked out (4). In an adequate distance from the hole edge the average value of the Vickers hardness is 150 HV 0.2 for the uninfluenced area. From 2.5 mm distance up to the hole edge we found an increase of the hardness up to a value of about 330 HV 0.2. Figure 5 shows the effect of cold-work hardening due to punching.

Figure 5. Distribution of Vickers hardness HV 0.2 in the nearer area around a punched hole.


With the cold-work hardening the tensile strength will also increase. To check this, we made tension tests with micro tensile specimen. The cross section of the test pieces is only 2.0 x 0.5 mm. So it is possible to produce five test specimen out of the close-up range of a punched hole. The position of the specimen is tangential to the edge of the hole. Figure 6 shows the results of the test in form of stress-strain curves. Here you can see the curve M5 which is typical for steel S 235 JR G2. This specimen has a distance of 2.7 mm measured from the hole edge and lies in the uninfluenced area. If you now look at the curves of the specimen which were nearer to the hole edge you can see that the tensile strength of the material increases and that the elongation at fracture decreases rapidly. This means a loss of ductility and an embrittlement due to cold-working during the punching process. This material data explains why there are so many small cracks on the surface around a punched hole, while a member with drilled holes has only one final crack in failure mode.

Figure 6. Stress-stain curves of the micro tensile test specimen

(x distance of the test piece from the edge of the hole). SLIP AND CREEP TESTS Many slip tests under static load have been carried out over the years. In the following some actual friction coefficients µ of double lap joints with different surface treatments are represented µ = 0.2 for mill scale, µ = 0.2 for hot dip galvanized surfaces with pure zinc on it, µ = 0.35 to 0.38 for hot dip galvanized surfaces without pure zinc,

µ = 0.6 for hot dip galvanized surfaces without the removed pure zinc, but with additional alkali-zinc-silicate friction paint.

Specimen with pure zinc on its surfaces were dipped into the zinc bath without removing the pure zinc layer afterwards, they had a glossy appearance. To separate and remove the pure zinc layer the members were put in a centrifuge after galvanizing. The appearance of these specimen was matte and lacklustre. Friction coefficients for different surface treatments which result from former slip tests can be find in the literature. As a result it can be seen that the friction paint produces a remarkable increase of the friction capacity. But it is necessary to check under which percentage of that value µ the connection does not slip-creep and maintains its rigidity under static long term loads. Therefore, we examined the creep behaviour of such connections (5).


Figure 7 presents three creep-slip tests of a high strength friction grip connection with hot dip galvanized contact surfaces with alkali-zinc-silicate friction paint. In theses cases the long term test loads were 90%, 68% and 50% of the friction capacity taken from short term tests. The 90%-test did not fulfil the requirement for displacement (< 300 µm within 30 years), while the 68%- and the 50%-tests did succeed. Of course, all curves in the Figure, which resulted from one year or two year tests, have been extrapolated in a log-linear diagram according to EC 3. This extrapolation is shown by the crosses.

Figure 7. Creep-slip curves of long term tests with alkali-silicate-zinc-paint on

the hot dip galvanized contact surfaces. After these tests the question came up as to whether the friction capacity µ may be influenced by a long term loading, such as included by the creep tests. Therefore, the behaviour of the slip connection with hot dip galvanized contact surfaces having alkali-zinc-silicate friction paint has been tested after creep tests, while in the creep tests the actual friction has been exploited only up to 60% till 70%. The subsequent slip tests were performed until slip occurred and they should give the remaining friction capacity. Figure 8 shows in the left diagram the results of the initial short term slip tests with HVM16 and HVM20 bolts. The series II and III in the middle of the diagram show the ultimate slip capacities after long term tests (no reduction!). Series IV at right hand in the diagram show the slip capacities after fatigue tests with two million load cycles (no reduction!).

Figure 8. Behaviour of the slip coefficient of hot dip galvanized surfaces with alkali-silicate-zinc-paint during long term testing and fatigue loading.


Besides the tests described above, another treatment of the contact surfaces has been investigated, namely high strength friction grip connection with hot dip galvanized contact surfaces with a slight shot blasting (sweeping) instead of alkali-zinc-silicate friction paint. The loading procedure was the same as before. The result of the tests was analogous. After long term tests (series II and III) and after fatigue tests with two million load cycles (series IV) there was no reduction of the ultimate slip capacities in either case. A last question: Is there constancy or an increase of the displacement during fatigue loading up to more than one million (M) cycles? The solution is illustrated by figure 9. Under high exploitation of about 70% of the static slip capacity and after 1.1M to 2M cycles the displacement ended in a maximum 0.12 mm. We see a kind of hysteresis which converges on a value much below the limit of 0.300 mm. Nine more tests show the same, or even better, results. Further investigation have been performed on hot dip galvanized steel members with punched slotted holes, friction paint and high strength preloaded bolts. Results can be taken from the literature (6).

Figure 9. Creep displacement for hot dip galvanized connections with

preloaded high strength bolts during fatigue loading. A frequently asked question is: Can we rely on a certain constancy of the bolt preload FV or do we need to take into account a considerable reduction? This takes place since in a steel package there are at least four galvanized zinc layers on the members, another four galvanized zinc layers on the washers and two layers on the bolt head and nut when clamped together. The total amount of zinc layer may add up to 0.5 mm, which surely creeps under the high preload of the bolt. Therefore, a second tightening procedure should go over the connections after two hours or an overtightening of the bolts should be envisaged. But what is the amount of the preload reduction? Since in our tests the preload is permanently measured, we can give relevant information on this topic. Apart from that there is the possibility of overtightening the bolts by a well defined percentage to cover the creep of the bolt force. Rotation-preload diagrams of six tightening tests with HV bolts M 20x100 - 10.9 show that an additional rotation of 20° makes an increase of about 10% of the required preload according to DIN 18800-7, EC 3 and EN 1090, which usually covers the creep influence.


FATIGUE TESTS Masts and towers are, in addition to their dead load, mainly loaded by wind. The wind may come from different directions, leading to oscillating movements and stresses in the structure and in the members as well as in the joints. The resulting number of cycles can be very high and is in the range of the fatigue strength. The use of bearing-type connections in hot dip galvanized members with punched holes leads to questions about the load carrying behaviour of these joints under cyclic loading. The knowledge of the fatigue behaviour of the joints is necessary to predict the lifetime of the structures. In earlier research projects at the TUHH, low-cycle-fatigue and high-cycle fatigue tests were carried out. Members with holes and bearing-type connections with both punched and drilled holes, but without any preload of the fasteners were examined (7, 8). The test specimens consisted of S 235 JR G2 (formerly: RSt 37-2) and the loading was of simple sinus wave form, while the ratio κ between the lower and upper tension in the net section was +0.1. The test specimen for the shear connection (type 1) and the member with a hole (type 2) are shown in figure 10. The bolts were hand tightened after they had contact with the wall of the hole in order to exclude the influence of friction due to preloading of the bolts. The dimensions of the test specimens were selected primarily to produce a net section fatigue failure of the middle member, and not a failure in the bolts or cover plates.

Figure 10. Test specimen of bearing-type connections (type 1) and members with a hole (type 2) for fatigue testing.

After the fatigue failure of the test specimens we examined the crack surfaces to get information about the crack initiation. Figure 11 shows two representative net section failures of members with a hole. It was remarkable that the starting point of the crack front for punched holes lies in the first zone on the upper corner of the hole edge (figure 11a) and never in the third zone where the punch leaves the plate. From here the fatigue crack front runs through the material until the net section is so much weaken that finally a static crack causes the rupture. In members with drilled holes a surface crack at the wall of hole is predominant (figure 11b).


Figure 11. Fatigue rupture mode with marked crack fronts after fatigue loading.

INFLUENCE OF PUNCHING AND GALVANIZING The analysis of the fatigue tests happens grafically as S-N curve (stress S over load cycles N) in log-log diagrams. The higher the curve in the diagram the higher is the fatigue resistance of the test specimen. The experimental S-N curves of different hot dip galvanized test specimens are shown in figure 12. Members with hole and bearing-type connections are compared. As expected, the members with a hole were able to withstand a higher stress range ∆σ at the same number of cycles N up to failure than the joints. A comparison between the test specimen with punched holes and the test specimen with drilled holes shows the negative influence of punching. The S-N curve for both different structural members with punched holes lies under the corresponding S-N curve for drilled holes.

Figure 12. S-N-curves of hot dip galvanized members with hole and shear-bearing connections with drilled and punched holes, FV = 0 (L = member with a hole, V = bearing-type connection s = punched hole, b = drilled hole, f = hot dip galvanized, κ = stress relation).

Fatigue tests of hot dip galvanized and non-galvanized bolted connections with punched and drilled holes were carried out to determine the influence of galvanization performed after punching or drilling. The results are compared in figure 13. It can be seen clearly that the fatigue life decreases owing to galvanizing and punching. This is the case for members with


both punched and drilled holes. The figure also shows that the influence of punching on fatigue behaviour is roughly equal to influence of galvanizing. This means that a non-galvanized structural member with a drilled hole has the highest fatigue resistance, for example 2M cycles at a constant stress range ∆σ = 80 N/mm2. If the member has a punched hole or is galvanized, the influence is nearly the same; the fatigue life decreases with a ratio of 2.0. Now the fatigue failure for a stress range ∆σ of 80 N/mm2 is at 1M cycles. If the member is both punched and galvanized there is an additional effect and the number of load cycles decreases to 500,000.

Figure 13. S-N-curves of non-galvanized and hot dip galvanized shear-bearing connections with drilled and punched holes, FV = 0 in comparison with EC 3 (nf = non galvanized).

COMPARISON WITH EUROCODE 3 Figure 13 shows a comparison of experimental results of tested bearing-type connections. They are classified as detail category 112 of EUROCODE 3. From N = 10,000 to a constant amplitude fatigue limit at N = 5 x 106 the slope of the curve is m = 3.0. The slope then changes to m = 5.0 up to the cut-off limit (N = 1 x 108). All results of the tests are clear below the S-N curve of EC 3. The S-N curve of EC 3 shows the material resistance with P = 95% confidence interval and 5% probability of failure, while the mean fatigue strength line of the tests has a P = 50%. Therefore, the statistical analysis confirms that the experimental results are even lower than the S-N curve of EC 3. The experimental tests for members with a hole have the same results. This leads to the conclusion that dimensioning using EC 3 is not safe for these structures. This is valid for hot dip galvanized members as well as for non-galvanized members with drilled or punched holes. The relation of these results to the corresponding S-N curves of the EC 3 is of great interest with regard to the position of the points of failure and the slope of the S-N curve relative to the S-N curves of EC 3. The slope of the fatigue strength curves of bolted joints lies between m = 5.9 and m = 6.7. The results show that the slope of the S-N curves with m = 3.0 is too high for these experimental curves. In literature (9) we found other fatigue tests of connections in comparison with EC 3. The result of the experimental tests is the same as before. Most of the fatigue life data lies below the S-N curve of EC 3. This leads to the question: Is the EC 3 is suitable for use? Because the detail category 112 is specifically permitted to non-preloaded high strength bolts.


THEORETICAL INVESTIGATIONS Nevertheless, independent from the comparison with EC 3 it is necessary to strengthen the bolted connections in hot dip galvanized steel constructions under fatigue loading. The reason for the low fatigue resistance of bearing-type connections is the unfavourable stress distribution in the net section of the member. Near to the hole there is a high peak of the notch stress. In the 1950s and 1960s high strength friction grip connections with preloaded bolts were developed (10). Tests have shown that the fatigue behaviour of members with preloaded bolts is much better than that with non-preloaded bolts. The increase in capacity, which means the increase in stress range ∆σ or load cycles N was so great that sometimes the fatigue behaviour of plain bars could be reached. This is due to the high pressure under the washers of the bolts. This high pressure gives a certain protection of the area around the hole, so that the stress distribution in the net section became much more favourable, even after slip of the connection. FATIGUE TESTS WITH PRELOADED FASTENERS The idea of covering the area around the hole by a high strength preloaded bolt can be transmitted to the problem of fatigue behaviour of hot dip galvanized steel members with punched holes. In a 2003 finished research project on this topic many fatigue tests were carried out at the TUHH. The aim was the strengthening of the fatigue resistance of the constructions. For this the negative influence of punching and galvanizing on material data near to the hole must be reduced. To ensure the comparison between the fatigue tests of the earlier research project of the members without any preload, we use the same geometry of the test specimen for this project (see figure 10). In the diagram of figure 14 the results of various fatigue tests can be seen. Hot dip galvanized bearing-type connections with punched holes were compared with friction-type connections having 50% and 100% preloaded high strength bolts and with an unnotched specimen. The open circles and the lowest S-N-curve show the load cycles of connections with non preloaded bolts. The full grey circles represent the results of equivalent connections with preloaded bolts tightened up to 50% of the required preload for a high strength bold M16. Even a preload of only 50% increases the fatigue behaviour enormously. The increase can be better expressed by the stress range ∆σnet than by the number of cycles. We found a step from 72 N/mm2 to about 180 N/mm2 at 1 million cycles. There was a further increase of fatigue strength up to 233 N/mm2 at a preload of 100% of the required preload (full black circles). This value is near to the fatigue failure of unnotched bars. They reach a stress range ∆σnet of 280 N/mm2 also regarded at 1 million cycles. These regarded S-N curves for the friction-type connections have a friction coefficient µ of 0.38 between the middle member and the cover plates. This value is valid for hot dip galvanized surfaces without any pre-treatment. As a consequence of the influence of friction identified previously, we painted the surface of the members with alkali-silicate coating to increase the friction coefficient. With this treatment we achieve a value of µ = 0.6. The results of the fatigue tests are also shown in figure 14. The open grey circles represent the results of tests with 50% preloaded bolts and the open black circles the results with 100% preload. It can be seen that it is possible to achieve a further increase of the fatigue resistance in comparison with the corresponding S-N curves for friction-type connections without any treatment of the hot dip galvanized surface.


Figure 14. Comparison between S-N curves for hot dip galvanized bearing-type

connections with punched holes, friction-type connections with 50% and 100% preloaded bolts and unnotched specimen (RS = reference tests with unnotched bars, asz. = alkali-silicate-zinc-coating).

In the next diagram (figure 15) we compared the S-N curves again with the corresponding S-N curves for non-galvanized connections with drilled holes and with the S-N curve of EUROCODE 3 (detail category 112). As seen before, the preload increases the fatigue behaviour of the connections. One interesting scientific finding is that the negative effect of punching in combination with galvanizing is neutralized because the experimental S-N curves for 50% and 100% are nearly identical. And these S-N-curves now lie above the S-N curve of the detail category 112 of EC 3 for high number of load cycles.

Figure 15. S-N-curves of non-galvanized and hot dip galvanized connections with drilled and punched holes with preloaded high strength bolts, FV = 0, 50% and 100% of 0,63⋅fu,b⋅ASp in comparison with EC 3 (detail category 112).

The problem which is still exists is the evaluation of the slope with m = 3.0. The real behaviour of the tests shows that the slope is smaller than m = 3.0. In particular, the preload causes a smaller slope, which can be explained by the crack growth. The crack can only be observed in the protected area under the washers. If the crack leaves this area the rupture of


the member begins directly. When the slope of the EC 3 is changed from m = 3.0 to m = 5.0 there is a better fitting for the whole S-N curve (see figure 15) and especially for the low-cycle-fatigue behaviour. The slope of m = 5.0 is chosen with reference to DS 804 which is valid for railway structures. In DIN 15018 and DIN 4132 the slope is once more flatter. Besides the fatigue resistance the crack growth is of great interest. The literature gives many information about the stress distribution of members with cracked and uncracked net sections but we found nothing about the influence of the pressure under the washers of a preloaded high strength bolt relating to this problem. Based on experimental tests and finite element parameter studies we did further investigations to clear the effects of the preload on the notch stresses near to the hole, the crack initiation and the crack growth. Finally a simplified calculation method for the lifetime prediction of the steel members was developed (11, 12). With that a large test programme with static, low-cycle- and high-cycle-fatigue tests is finished which had the aim to give answer about the bearing capacity and the fatigue behaviour of hot dip galvanized steel members with punched holes. CONCLUSION The idea of protecting the net section area around a punched hole of a hot dip galvanized member by the use of preloaded high strength bolts with two washers has shown that a remarkable influence on the fatigue life can be achieved. The advantage of this method is the ease of handling with maximum of efficiency. Even the negative influence of punching in combination with hot dip galvanizing can be neutralized. ACKNOWLEDGEMENTS We greatly acknowledge that the institution AiF (Arbeitsgemeinschaft industrieller Forschungsvereinigungen in Bonn, Germany), the Minister of Economy of the Federal Republic of Germany, Berlin, and the GAV (Gemeinschaftsausschuss Verzinken e.V., Düsseldorf) gave financial supported this research (Research project AiF no. 11097/N1 and AiF no. 12547/N1). REFERENCES (1) Valtinat, G., Dangelmaier, P. (1993). Schraubenverbindungen mit gestanzten Löchern

in zugbeanspruchten, feuerverzinkten Bauteilen. Bericht Nr. 119 des Gemeinschaftsausschuß Verzinken e.V., Forschungvorhaben GAV-Nr. FD 18, AiF-Nr. 7448, Düsseldorf.

(2) Valtinat, G., Wilhelm, M. (1995). Vergleich des Last-Verschiebungs-Verhaltens und der Traglast von Schraubenverbindungen mit gestanzten und gebohrten Löchern in zugbeanspruchten, feuerverzinkten Bauteilen. Bericht Nr. 129 des Gemeinschaftsausschuß Verzinken e.V., Forschungsvorhaben GAV-Nr. FD 20, AiF-Nr. 9305, Düsseldorf.

(3) Valtinat, G., Grycz, J., Wilhelm, M. (1991). Feuerverzinkte Stahlbau-Verbindungen mit hochfesten Schrauben und gestanzten Löchern. Vortrags- und Diskussionsveranstaltung 1990 des GAV, Düsseldorf, Germany, ISSN 0344-3582, 115-132.

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vorgespannten Schrauben. Gemeinschaftsausschuß Verzinken e.V., Forschungsvorhaben GAV-Nr. FD 23/II, AiF-Nr. 12547/N1, Düsseldorf.

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(8) Valtinat, G., Huhn, H. (2000). Betriebsfestigkeit von stählernen Lochstäben und Schraubenverbindungen mit feuerverzinkten Bauteilen und gestanzten Löchern. Gemeinschaftsausschuß Verzinken e.V., Forschungsvorhaben GAV-Nr. FD 23, AiF-Nr. 11097/N1, Düsseldorf.

(9) DiBattista, J. D., Adamson, D. E. J., Kulak G. L. (1998). Fatigue Strength of Riveted Connections. Journal of Structural Engineering, 124 (7), 792-797.

(10) Steinhardt, O., Möhler, K. (1954). Versuche zur Anwendung vorgespannter Schrauben im Stahlbau, I. Teil. Berichte des Deutschen Ausschusses für Stahlbau, Heft 18.

(11) Hadrych, I. (2000). Wachstum von Ermüdungsrissen an Niet- und Schraubenlöchern unter Berücksichtigung von Vorspannkräften der Verbindungsmittel. Dissertation, Technische Universität Hamburg-Harburg.

(12) Huhn, H. (probably published 2004). Ermüdungsfestigkeit von Schraubenverbindungen aus feuerverzinkten Stahlbauteilen mit gestanzten Löchern. Dissertation, Technische Universität Hamburg-Harburg.


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